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Experimental investigation on thermal management performance of an integrated heat sink with a piezoelectric micropump
Applied Thermal Engineering 161 (2019) 114053
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Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
Research Paper
Experimental investigation on thermal management performance of an integrated heat sink with a piezoelectric micropump
T
Yong Tanga, Mingze Jiaa, Xinrui Dinga, , Zongtao Lia, Zhenping Wana, Qinghong Lina, Ting Fub ⁎
a b
Key Laboratory of Surface Functional Structure Manufacturing of Guangdong High Education Institutes, South China University of Technology, Guangdong, China Key Laboratory of Metallurgical Equipment and Control Technology, Wuhan University of Science and Technology, Wuhan 430081, China
ARTICLE INFO
ABSTRACT
Keywords: Heat sink Piezoelectric micropump Thermal management
High integration density is the trend of future electronics. The accompanying heat dissipation remains a serious problem in electronic thermal management. To overcome the high power density of heat dissipation, an integrated heat sink with a piezoelectric micropump (IHS-PMP) was developed. The driving performance of the IHS-PMP, including the output flow rate and pump pressure, was established to investigate the effects of the driving voltage and driving frequency. Additionally, it was found that the working fluid generated bubbles under a thermal load. The inevitable bubbles significantly affected the flow rate of the IHS-PMP. Finally, the effects of the heating power and driving frequency on the thermal management performance of the IHS-PMP were experimentally investigated and analyzed. The results indicated that with the optimal driving parameters of the IHS-PMP and at a heating power of 80 W, the wall temperature rise was 47.3 °C. This indicates that the IHS-PMP had excellent thermal management performance. The IHS-PMP can be regarded as a high-potential component in loop liquid cooling, which is beneficial to the thermal management of electronic devices.
1. Introduction Thermal management has become a major challenge in the development of electronic components owing to the increased heat flux and reduced component size [1–4]. The reliability of electronic components and optoelectronic materials decreases dramatically as the temperature of the component increases [5–8]. The failure probability of an electronic component increases exponentially when the temperature is greater than 75 °C [9]. According to statistics, more than 50% of electronic component failures are caused by a thermal issue [10]. Additionally, the installation space for thermal management components has become increasingly narrow. The effective thermal management and heat dissipation of electronic devices in a small space has become a bottleneck, restricting the development of electronic devices [11–13]. Loop liquid cooling technology has received extensive attention owing to its stable and efficient long-distance heat transfer, and flexible layout. Pal et al. [14] proposed a liquid cooling system for aircraft power electronic systems and discussed the system-level issues for liquid cooling. Kang et al. [15] introduced the design and properties of an advanced liquid cooling system that can cool single or multiple heat sources in a computer system. The cooling system employs copper cold plates with medium-scale channels to extract heat from the central processing unit and graphics processing unit heat sources, using a ⁎
highly efficient liquid-air heat exchanger with flat copper tubes and common fins to transfer heat to the air via forced convection. A highly reliable and compact pump was used to circulate the fluid in a closed loop. The entire system has a very low fluid permeability and a long service life. However, the loop liquid cooling system contains many components, such as drive pumps, heat sinks, radiators, and pipes. These large components occupy too much cooling space to be used in electronic devices that operate in small spaces. Currently, the driving components of the loop liquid cooling system are predominantly rotor pumps [16,17], which have a loud operating noise, large size, and are not suitable for use in compact loop liquid cooling. Among the numerous micro-driven pumps, the driving abilities of the electrohydrodynamic micropump [18] and the magnetohydrodynamic micropump [19] are weak. The heat-driven micropump [20] is significantly affected by the temperature. The electrostatic actuation micropump [21,22] is difficult to manufacture owing to the use of silicon micromachining and the high precision of the distance between the diaphragm and the electrode. However, the piezoelectric micropump has the advantages of a small temperature influence and a strong driving capability and is very suitable for application in micro-loop liquid cooling systems. Researchers have studied the piezoelectric micropumps used in the loop liquid cooling system [23–28]. Ma et al. [23] replaced the conventional drive
Corresponding author at: South China University of Technology, Guangdong 510641, China. E-mail address: [email protected] (X. Ding).
https://doi.org/10.1016/j.applthermaleng.2019.114053 Received 13 March 2019; Received in revised form 12 June 2019; Accepted 28 June 2019 Available online 29 June 2019 1359-4311/ © 2019 Elsevier Ltd. All rights reserved.
Applied Thermal Engineering 161 (2019) 114053
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Nomenclature IHS-PMP PVPC HTC Tin Tout Tci Tw Q ΔV f ξ U ΔV′ P qt qe ql qe′ Ap Rt Rc Rs λc
λs lc
integrated heat sink with a piezoelectric micropump pumping volume per cycle, mL heat transfer coefficient, W m−2 °C−1 inlet fluid temperature, °C outlet fluid temperature, °C thermocouple reading, °C wall temperature, °C theoretical output flow rate without thermal load, mL/min the amount of change in chamber volume due to vibration of the piezoelectric disc in a single cycle, mL frequency of sinusoidal AC power supply, Hz piezoelectric drive efficiency coefficient driving voltage, V pumping volume per cycle, mL pump pressure, kPa total heating power, W effective heating power, W loss of heating power, W effective heat flux supplied to test section, kW/m2 The ratio of the loss of heating power to the total heating power area of the heat exchange contact surface, mm2 total thermal conduction resistance, °C/m thermal conduction resistance of pure copper, °C/m thermal conduction resistance of solder, °C/m thermal conductivity of pure copper, W m−1 °C−1
ls ΔT Ta R’ h Q’ c ρ
thermal conductivity of solder, W m−1 °C−1 distance between the upper surface of the copper block and the thermocouple, mm thickness of the solder layer, mm wall temperature rise, °C average temperature of the inlet and outlet liquid temperature, °C thermal resistance of the IHS-PMP, °C/W heat transfer coefficient, W m−2 °C−1 flow rate of IHS-PMP in the thermal management test, mL/min specific heat capacity of deionized water, J kg−1 °C−1 density of deionized water, g/cm3
Subscript in out ci w t e l p c s a
pump with a piezoelectric micropump, which was connected to a heat sink to form a new liquid cooling system. It satisfies the system miniaturization requirements and achieves good thermal management performance. Researchers have improved the structure of piezoelectric micropumps. Ma et al. [24] improved the full-side support mode of the piezoelectric vibrator and proposed a single-sided support cantilever assembly method, which can increase the effect of the piezoelectric vibrator on the volume of the pump chamber. The specific effect is to increase the volume change. Shabanian et al. [25] improved the
inlet outlet thermocouple location in stream-wise direction wall total effective loss platform of on the upper surface of the copper block copper solder average
structure of the piezoelectric vibrator, proposing a ring-shaped structure, which improved the displacement and the output performance of the piezoelectric micropump. Ren et al. [26] improved the check valve structure and proposed a long diaphragm-shaped check valve, which can significantly improve the operating frequency of the piezoelectric pump. Additionally, other researchers have proposed miniaturizing the heat sink in the liquid cooling of the loop to further reduce the size of the radiator [27,28]. The aforementioned studies were aimed at miniaturizing the entire
Fig. 1. Diagram of IHS-PMP ((a) is the structure diagram of IHS-PMP. (b) is the structure diagram of the heat exchange chamber). 2
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coolant (deionized water) to flow uniaxially through the heat exchange chamber. The drive unit was composed of a piezoelectric disc and a double O-ring. The piezoelectric disc (P-58, peak-to-peak value of the maximum allowable driving voltage was 300 V) was composed of a piezoelectric ceramic and a brass foil bonded together. The piezoelectric ceramic had a diameter of 36 mm and a thickness of 0.38 mm. The brass foil had a diameter of 41 mm and a thickness of 0.18 mm. One electrode of the piezoelectric disc was soldered to the piezoelectric ceramic, and the other was soldered to the brass foil. In this study, it was necessary that the piezoelectric disc operated in a heated state, and the sealing property was more demanding. Therefore, a double O-ring sealing element structure was designed. The double O-ring was similar to the overlap of two O-rings, except that there was a 0.2-mm-thick connection between the two O-rings, improving the sealing performance. The piezoelectric disc was clamped by a double O-ring with a diameter of 42 mm to form the drive unit. A pressing plate with a thickness of 3 mm was used to compress the double O-ring of the drive unit. The stainless-steel gasket was used to limit the amount of depression of the pressing plate.
heat dissipation system by reducing the size of various components in the loop liquid cooling. However, complex system components in the loop liquid cooling system lead to increased connection points of various components, which may lead to leakage of the working fluid, resulting in system failure or even damage to the microelectronic devices. Therefore, the development of new heat dissipation components and the improvement of the integration of the loop liquid cooling system to satisfy the high heat flux thermal management requirements in the confined spaces of electronic devices are urgently needed. Integration technology reduces the number of components and realizes the multifunctional application of components [29–31]. The integration of a piezoelectric micropump and a micro-heat sink is a feasible solution for effective thermal management. At present, the radiator is integrated with the vapor chamber [32–34] and heat pipe [35,36], which can make the radiator structure compact, reduce the thermal resistance, and improve the thermal management performance. However, there has been little research on the integration of an efficient piezoelectric pump with a micro heat sink. In this work, a novel integrated heat sink with a piezoelectric micropump (IHS-PMP) is proposed to satisfy the high heat flux thermal management requirements in the confined spaces of electronic devices. The IHS-PMP driving performance without a thermal load and the thermal management performance were experimentally evaluated.
3. Experimental setup and procedure In this work, two measurements were performed to synthetically investigate the performance of the IHS-PMP, the driving performance without a thermal load (including the output flow rate measurement and pump pressure measurement), and the thermal management performance (including the temperature rising and thermal resistance measurements). The device schematic is shown in Fig. 2. The power drive part include an arbitrary waveform generator (AFG3251, maximum output voltage amplitude of 5 V) and a power amplifier (ATA2210, maximum magnification of 100). Deionized water was utilized as the working liquid. The deionized water was in a constant temperature water tank with an automatic temperature regulation system. The temperature of the constant temperature water tank was set as 25 °C. The driving performance test platform is shown in Fig. 2(a). As the
2. Design and fabrication of IHS-PMP A diagram of the IHS-PMP is shown in Fig. 1. The IHS-PMP comprised four parts: a heat exchange chamber, a drive unit, a stainless steel gasket, and a pressing plate. The copper heat exchange chamber was fabricated by milling a cylinder with a depth of 4.8 mm and a diameter of 33 mm. The design principle of the IHS-PMP heat exchange chamber was to use a part of the piezoelectric pump as a heat sink; that is, the heat sink and the piezoelectric pump shared a chamber for integration purposes. Thus, the IHS-PMP had the functions of pumping liquid and heat transfer. Two check valves were utilized to cause the
Fig. 2. Diagram of experimental setup ((a) is the flow rate and pump pressure test device diagram of IHS-PMP without thermal load. (b) is the thermal management performance test device diagram of IHS-PMP). 3
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series flow meter affected the pump pressure of the system, it also further affect the flow rate. The flow rate test section comprised a beaker, an electronic balance, and a timer. The flow rate of the IHSPMP without a thermal load was measured by averaging the pumped water mass within 1 min. Two lifting platforms were used to adjust the water tank height and output piping to ensure that the water tank, output piping, and IHS-PMP had the same height. The pressure test part comprised a differential pressure transmitter, a direct-current (DC) stabilized power source, a pressure relief valve, and a temperature acquisition card. The outlet of the IHS-PMP was connected to the differential pressure transmitter (PCM300, range of 0–20 kPa) and the pressure relief valve via a three-way valve; the pressure relief valve was used to vent the pressure in the tube after the test was completed. Similarly, the tank, IHS-PMP, and pressure transmitter were at the same height. The DC stabilized power source supplied power to the differential pressure transmitter and the temperature acquisition card. Once the pressure was steady, it was recorded by the pressure transmitter with a sampling rate of 1 Hz for a period of 1 min. A diagram of the thermal management performance test part is shown in Fig. 2(b). The constant-temperature (25 °C) deionized water in the water tank was circulated by the HIS-PMP. A cross-sectional view of the thermal management performance test device is shown in Fig. 3. Nine heaters were embedded inside a copper cylinder with a diameter of 33 mm. The test sample was soldered on top of the copper block using a thin layer of solder (Pb-Sn-Ag-Sb) to minimize the contact thermal resistance. The nine heaters were powered by a variac and a wattmeter showed the input power. The heating area was embedded into a thermal insulation material (thermal conductivity of ≤0.27 W/ m °C) to reduce the heat loss. In the data acquisition part, seven K-type thermocouples were employed to measure the temperature, with two at the inlet/outlet of the IHS-PMP (Tin, Tout), and three (T1, T2, T3,) were distributed 5 mm below the test surface with intervals of 12 mm. In the vertical direction, two K-type thermocouples (T4, T5) were placed below T2 with an interval of 6 mm. A HIOKI data acquisition system (LR841030 wireless logging station and LR8510 wireless voltage/temp unit) was employed to collect and display the temperature using the seven K-type thermocouples. The wall temperature (Tw) was then calculated from the Fourier Law.
4. Data reduction 4.1. Theoretical output flow rate and output pressure of IHS-PMP The design of the IHS-PMP is based on the principle of a piezoelectric micropump and its output flow rate can be calculated as follows: (1)
Q = 60 Vf
where ΔV is the amount of change in the chamber volume due to the vibration of the piezoelectric disc in a single cycle, f is the frequency of the sinusoidal alternating-current power supply, and ξ is the piezoelectric drive efficiency coefficient. Stemme et al. [37,38] proposed a method for calculating ΔV. The relationship between ΔV and U is:
V
(2)
U
where U is the driving voltage applied to the piezoelectric disc. To facilitate the study of the relationship between the driving frequency and the output flow rate, the pumping volume per cycle (PVPC) of the IHS-MPM is defined as the amount of liquid volume pumped by the IHS-MPM when the piezoelectric disc vibrates for one cycle and is denoted as ΔV′:
V' = V =
Q 60f
(3)
The working fluid used in this experiment was deionized water, which can be considered as an incompressible fluid. According to the assumption that the piezoelectric drive pump chamber is a closed space [37,38], the relationship between the pump pressure P and U is:
P
U
(4)
4.2. Wall temperature of IHS-PMP is:
The total heating power qt in the thermal management experiment
qt = qe + ql
(5)
where qe is the effective heating power applied to the heated surface
Fig. 3. Details of the thermal management performance test device (Rc is the thermal conduction resistance of the copper block, Rs is the thermal conduction resistance of the solder layer, Rt is the total thermal conduction resistance, R' is the thermal resistance of the IHS-PMP). 4
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and ql is the loss of heating power. An insulating cotton with a thickness of 20 mm was wrapped around the outside of the actual test platform. The one-dimensional Fourier Law was employed to calculate the actual heat flux densityqe' according to T2, T4, and T5:
qe' =
c
ql = qt
dT dx
(6)
qe' Ap
(7)
h=
ql qt
Q' =
qe
(9)
qe (10)
Ap
Because the direction of the heat flux flow was straight upward, the wall temperature was defined as the surface temperature of the IHSPMP, which was the temperature of the upper surface of the solder. Generally, in the IHS-PMP chamber, the temperature of the fluid increased linearly along the flow direction. According to the one-dimensional heat conduction equation, we define the temperature at the midpoint of the IHS-PMP as the wall temperature Tw:
Tw = T2
The driving frequency and driving voltage significantly affected the output flow rate and pump pressure of the piezoelectric micropump. The output flow rate of the IHS-PMP had a large influence on the thermal management performance. Therefore, before the thermal management performance test of the IHS-PMP, it was necessary to study its driving performance without a thermal load, focusing on the effects of the driving voltage and driving frequency on the output flow rate of the IHS-PMP. The thermal management performance test of the IHS-PMP was based on the results of the driving performance without a thermal load.
where Rt represents the total thermal conduction resistance along the heat conduction direction, which can be calculated as follows: (12)
Rt = R c + Rs
where Rc represents the thermal conduction resistance of the copper block, and Rs represents the thermal conduction resistance of the solder layer:
Rc =
lc , Rs = c Ap
ls s Ap
5.1. Driving performance of IHS-PMP without thermal load The driving voltages of 200 and 300 V were selected to study the influence of the driving frequency on the driving performance of the IHS-PMP. The check valve response frequency is generally within 100 Hz; thus, the driving frequency was optimized in the frequency range of 0–100 Hz. The driving frequency parameters were set in the range of 0–100 Hz, with intervals of 5 Hz. Fig. 4 shows the output flow rate of the IHS-PMP. The maximum output flow rates for 200 and 300 V (140 and 210 mL/min, respectively) occurred at the frequency of 15 Hz. In the driving frequency range of 5–10 Hz, the piezoelectric disc vibrated at a lower frequency. As the driving frequency increased, the check valve gradually opened and the output flow rate increased. This is illustrated in Fig. 5; as the frequency increased, the PVPC increased. The PVPC reached its maximum value, and the check valve was fully open at 10 Hz. In the driving frequency range of 10–15 Hz, because the motion of the check valve lagged behind the vibration of the piezoelectric disc, the PVPC decreased. However, as the vibration frequency of the piezoelectric disc increased, the output flow rate of the IHS-PMP slowly increased. The maximum output flow rate was reached at 15 Hz. At 15–25 Hz, the PVPC decreased sharply owing to the lagging motion of the check valve behind the vibration of the piezoelectric disc. Even though the increasing driving frequency could increase the total number of pumping cycles, it was insufficient to compensate for the reduction in the PVPC. The output flow rate was gradually reduced at 15–25 Hz. When the driving frequency increased to 25–70 Hz, the second maximum of the output flow rate occurred in the driving frequency
(13)
Here, lc = 5 mm represents the distance between the upper surface of the copper block and the thermocouple (T1, T2, or T3), λc = 397 W/ (m °C) represents the thermal conductivity of pure copper, ls = 0.2 mm represents the thickness of the solder layer, and λs = 50 W/(m·°C) represents the thermal conductivity of the solder. The wall temperature rise is defined as:
T = Tw
Tin
(14)
where Tin represents the temperature of the thermocouple at the entrance. 4.3. Heat transfer coefficient (HTC) of IHS-PMP The average of the inlet and outlet liquid temperatures was taken as the fluid temperature:
Ta =
Tin + Tout 2
(15)
where Tout represents the temperature of the thermocouple at the outlet. The thermal resistance of the IHS-PMP can be calculated as follows:
R' =
(Tw
Ta ) qe
(18)
5. Results and discussion
(11)
qe Rt
Tin)
The dimensional errors due to the limitation of the milling accuracy and the solder layer thickness were estimated to be ± 0.01 mm. The accuracy of the temperature measurement by the precise K-type thermocouples was ± 0.2 K; The uncertainty of the power input was ± 1 W; The uncertainty of the pressure transmitter was ± 12 Pa; The uncertainty of the electronic balance was ± 0.01 g; Using a standard error analysis, the calculated uncertainties for the heat flux, the HTC, and the flow rate in the thermal management performance test were within 1.7%, 2.6%, and 1.6%, respectively [39].
The effective heat flux is:
qe' =
qe c (Tout
4.4. Uncertainties
in this experiment was determined to
qt
(17)
Ta
where c = 4186 J/(kg °C) represents the specific heat capacity of deionized water, and ρ = 1.0 g/cm3 represents the density of deionized water.
(8)
After testing, the be < 5%.Thus, ql ≪ 8qt.
Tw
To avoid the impact of the serial flowmeter on the IHS-PMP pressure and flow rate, in the thermal management test, the flow rate of the IHSPMP was calculated as follows:
where Ap is the area of the heat exchange contact surface (33 mm × 33 mm). The ratio of the loss of heating power to the total heating power is:
=
qe'
(16)
The HTC is calculated as follows: 5
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15, 20, 40, 45, and 50 Hz. According to the foregoing discussion, the two frequencies ƒ1 (15 Hz) and ƒ2 (45 Hz) exhibited reasonable output flow rates. Therefore, ƒ1 (15 Hz) and ƒ2 (45 Hz) were selected to investigate the influence of the driving voltage. The operating driving voltage ranged from 100 to 300 V, with intervals of 50 V. The IHS-PMP output flow rate at both driving frequencies increased linearly with the increasing driving voltage, as shown in Fig. 7. As the driving voltage increased, the deformation of the piezoelectric disc and the cavity change volume ΔV increased linearly, resulting in a linear increase in the output flow rate of the IHS-PMP. This is theoretically consistent with the relationship between the driving voltage and the output flow rate [38]. However, the driving voltage cannot be increased indefinitely. The maximum operating voltage of the piezoelectric disc selected in this study was 300 V. The IHS-PMP achieved the maximum output flow rate at 300 V. Additionally, the IHS-PMP pump pressure at both driving frequencies increased linearly with the increasing driving voltage, as shown in Fig. 7. This is theoretically consistent with the relationship between the driving voltage and the pump pressure [38]. Similarly, owing to the limitation of the operating voltage of the piezoelectric disc, the IHSPMP achieved the maximum pump pressure at a driving voltage of 300 V in this study.
Fig. 4. Flow rate and driving frequency relationship at different driving voltages.
5.2. Thermal management performance of IHS-PMP The thermal management performance of the IHS-PMP was tested under different driving frequencies. Six driving frequencies were selected with relatively high output flow rates, as previously mentioned. The driving voltage of 300 V with the maximum output flow rate was selected. The output heating power was increased from 0 W with intervals of 10 W. When the temperature test points T1–T5 fluctuated less than 0.5 °C within 1 min, the IHS-PMP was considered to be in a steady state. The upper limit working temperature was 70 °C. When the temperature test points T1-T5 of the IHS-PMP were > 70 °C, the IHS-PMP was considered to have reached the working limit, and the experiment was terminated. When the temperature of the IHS-PMP reached the upper limit working temperature, the heating power was defined as the upper limit heating power. If the IHS-PMP temperature fluctuated greatly and could not be stabilized within 15 min, the IHS-PMP was considered to have reached the stable limit working temperature and stopped heating. When the temperature of the IHS-PMP reached the limit working temperature, the heating power was defined as the stable limit working power. Fig. 8 shows the wall temperature rise of the IHS-PMP at the driving frequencies of 10, 15, 20, 40, 45, and 50 Hz. The wall temperature rise
Fig. 5. PVPC and driving frequency relationship at different driving voltages.
range of 45 Hz. When the driving frequency was 45 Hz, for the driving voltages of 200 and 300 V, the output flow rates of the IHS-PMP were 125 and 180 mL/min, respectively. In the driving-frequency range of 25–45 Hz, as the driving frequency increased, the phase lag of the check valve with respect to the piezoelectric disc exceeded one cycle; thus, the check valve was matched with the next cycle of the piezoelectric disc, and the output flow rate started to increase again. According to a previous study on the existing piezoelectric micropump [40], in the driving-frequency range of 45–70 Hz, the check valve cannot be fully opened, and the output flow rate decreases. Moreover, as the driving frequency increases, the minimum distance between the valve and the valve seat increases, and the phenomenon of incomplete closing is more pronounced, which results in a decrease in the output flow rate. Therefore, in the driving-frequency range of 25–70 Hz, a peak of the flow rate occurred when the driving frequency was 45 Hz. In the frequency range of 70–100 Hz, although the flow rate increased slightly at 80 Hz, the output flow rate was relatively low, the vibration frequency of the piezoelectric disc was too high, and the selected check valve did not effectively match the work. Therefore, this driving frequency range cannot be used for IHS-PMP operation. Fig. 6 shows the output pump pressure of the IHS-PMP in the driving frequency range of 5–100 Hz. The output pump pressure remained approximately constant at a high level. This indicates that the driving frequency had little effect on the IHS-PMP pump pressure. Therefore, the appropriate driving frequencies for the IHS-PMP samples were 10,
Fig. 6. Pump pressure and driving frequency relationship at different driving voltages. 6
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decreased as the heat power increased. At the driving frequency of 10 Hz, the flow rate decreased from 167.01 to 61.48 mL/min as the heating power increased from 10 to 80 W. Even the flow rate at 10 W was lower than the flow rate without heating. The decrease in the flow rate was caused by the generation of bubbles in the system, even though the water in the circulating water tank was degassed. Traditional thermal management performance tests of a heat sink are only executed at a certain flow rate. When the flow rate is chosen, it will not change with the increasing heat power. The actuating device in the loop system increases its power to maintain a stable flow rate when the heat power is increased. In this study, the thermal management performance tests of the IHS-PMP were performed under a constant actuating power. Even at the maximum heating power, the water temperature was < 73 °C; thus, the bubbles were not generated by the phase change. The thermal management performance test in this study was conducted in an open environment, which simulated the scene of the heat sink in the actual work. Before the test, the deionized water in the constanttemperature water tank was boiled sufficiently to remove non-condensable gases. However, as the test was conducted in an open environment, the environment of the constant temperature water tank could not be a vacuum; thus, some non-condensable gases were still dissolved in the water. When deionized water passed through the heat exchange chamber, the temperature difference between the deionized water and the wall changed the solubility of the gas in the water. Therefore, the gas precipitated out of the fluid, increasing the fluid resistance. This phenomenon occurs in a loop liquid cooling system operating in a non-vacuum environment [45,46]. The bubbles had a certain incompressibility which reduced the volume of the micropump chamber when the piezoelectric disc vibrated. This reduced the output flow rate of the IHS-PMP. Additionally, as the heating power increased, the temperature difference between the deionized water and the wall increased, which generated more bubbles, reducing the flow rate. Thus, the IHS-PMP followed the same principle (of flow-rate reduction), as the traditional piezoelectric micropump chamber [24]. This phenomenon is not caused by integration; it also occurs if the driving pump and the heat sink work in series. In the thermal management performance test of a traditional cooling system, the power of the driving pump is always increased to maintain the system output flow rate stability, but in this study, the driving part was integrated with the heat transfer part; thus, the driving parameters were not changed in the thermal management performance test. The IHS-PMP has a different ability to discharge air bubbles from the heat exchange chamber at different driving frequencies. When the driving frequency was 10 Hz, the pumping amount and output flow rate were maximized, and the bubbles were effectively discharged in the heat exchange chamber. Additionally, it was observed that the bubbles
Fig. 7. Effect of the driving voltage on the flow rate and pump pressure of the IHS-PMP.
Fig. 8. ΔT and heating power relationship at different driving frequencies.
was smallest in the driving frequency range of 10 Hz. As the driving frequency increased, the wall temperature rise also increased, but the wall temperature rise was greatest at 20 Hz. For instance, when the heating power was 40 W and the driving frequency was 10, 15, 20, 40, 45, and 50 Hz, the wall temperature rise of the IHS-PMP was 14.75, 16.73, 31.04, 20.66, 21.65, and 27.11 °C, respectively. Additionally, the ultimate power at different driving frequencies is shown in Fig. 9. The ultimate power of the IHS-PMP at 10, 15, 20, 40, 45, and 50 Hz was 80, 60, 40, 50, 50, and 40 W, respectively. The ultimate power of the IHSPMP at 10 Hz was limited by the upper limit working temperature, and at other driving frequencies, it was limited by the heat transfer stabilization before the test temperature reached the stable limit working temperature. The most suitable driving frequency for the IHS-PMP was 10 Hz. Compared with the thermal management method using aircooling [41] and phase-change heat transfer [32,42], under the same heating conditions, the thermal management performance of the IHSPMP in the loop liquid cooling was excellent at the driving frequency of 10 Hz. When the driving frequency was 10 Hz, compared with other thermal management methods using loop liquid cooling [43,44], the thermal management performance of the IHS-PMP was excellent at the same volume of the heat sink. The differences in the thermal management performances of the IHS-PMP were caused by the different driving frequencies, because at different driving frequencies, as the thermal power increased, the flow rate changed differently. Fig. 10 shows the variation of the IHS-PMP flow rate with the increasing thermal power at different driving frequencies. The flow rate
Fig. 9. Ultimate power at different driving frequencies. 7
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Fig. 12. HTC and heating power relationship at different driving frequencies.
Fig. 10. Flow rate and heating power relationship at different driving frequencies.
stably discharged. This is because at 10 Hz, the IHS-PMP had the lowest vibration frequency, which improved the stability of degassing. Although the driving frequencies of 15, 20, 40, 45, and 50 Hz yielded a considerable flow rate without a thermal load, the ability of the IHSPMP to discharge air bubbles was poor owing to the small PVPC. At the driving frequencies of 15 and 40 Hz, the bubbles oscillated back and forth at the outlet pipe. At the driving frequencies of 20, 45, and 50 Hz, as the heating power increased, the bubbles became larger and increased in number but did not move. The bubble movements in the chamber were considered to be similar. Bubbles that could not be discharged from IHS-PMP accumulated in the heat-exchange chamber. Therefore, at the same power, the amount of residual bubbles in the chamber was minimized when the IHS-PMP was at 10 Hz, the output flow rate was maximized, and the wall temperature rise was minimized. Because of the bubble generation and the decrease in the output flow rate, the thermal resistance of the IHS-PMP increased with the increasing heat power, as shown in Fig. 11. Compared with other driving frequencies, the thermal resistance was the lowest at 10 Hz. The IHSPMP was in the most efficient heat transfer state at 10 Hz. However, the overall thermal resistance of IHS-PMP is still relatively high, because inside the IHS-PMP, the water flow is considered to be laminar, and the flow velocity at the bottom of the heat exchange chamber is small. This part of the water directly exchanges heat with the heat exchange chamber, which will result in a relatively large thermal resistance of the IHS-PMP.
The HTC of the IHS-PMP is shown in Fig. 12. The IHS-PMP had the largest HTC at the driving frequency of 10 Hz, because the maximum output flow rate was 10 Hz with the minimal thermal resistance. Moreover, as the heating power increased, the HTC gradually decreased, because the flow rate gradually decreased. Additionally, 10 Hz yielded the optimum flow-rate stability, which is advantageous for heat exchange. Therefore, 10 Hz is the optimal driving frequency for heat exchange. At other driving frequencies, the output flow rate of the IHSPMP was low, and the liquid flow was unstable. This explains why the temperature of the IHS-PMP at other driving frequencies could not reach the upper limit working temperature. 6. Conclusion An IHS-PMP was proposed to address the thermal management demand of electronic devices in narrow spaces. The IHS-PMP solved the problems of large operating noise and a large rotor pump in loop liquid cooling. [16,17] Additionally, the problem of the leakage of working fluid caused by the piezoelectric micropump and the heat sink in series was solved by integration. [23–28]. The compactness of the loop liquid cooling system was improved via integration of a pump and a heat sink. First, the driving performance of the IHS-PMP without a thermal load was investigated. Then, the thermal management performance of the IHS-PMP was evaluated. The following conclusions were drawn.
• The output flow rate of the IHS-PMP without a thermal load was • •
significantly affected by the driving frequency owing to the influence of the matching of the piezoelectric disc and the check valve. The output pressure of the IHS-PMP without a thermal load was essentially unaffected by the driving frequency. The output flow rate and output pressure of the IHS-PMP without a thermal load were proportional to the driving voltage. The thermal management performance of the IHS-PMP was affected by not only the output flow rate but also the generation and discharge of bubbles in the heat exchange chamber. A high output flow rate and effective bubble discharge contributed to the thermal management performance of the IHS-PMP. The driving frequency of 10 Hz and driving voltage of 300 V were the best thermal management driving parameters for the IHS-PMP. Under these conditions, when the heating power of the IHS-PMP was 80 W, the wall temperature rise was 47.3 °C, which is suitable for the thermal management of electronic devices in narrow spaces.
Although the working mechanism and thermal management performance of IHS-PMPs have been studied, it is necessary to study the
Fig. 11. R and heating power relationship at different driving frequencies. 8
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structure of the heat exchange chamber and the bubble dynamics in the liquid to improve the heat management effect.
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Acknowledgements This work was supported by the National Natural Science Foundation of China (No. 51805173, No. 51735004), Science and Technology Program of Guangzhou, China (No. 201904010252), Fundamental Research Funds for the Central Universities (No. 2018MS44), The Open Fund of the Key Laboratory for Metallurgical Equipment and Control of Ministry of Education in Wuhan University of Science and Technology (No. 2018B01), Science and Technology Planning Project of Guangdong Province, China (No. 2015B020238004). Appendix A. Supplementary material Supplementary data to this article can be found online at https:// doi.org/10.1016/j.applthermaleng.2019.114053. References [1] T.C. Chang, S. Lee, Y.K. Fuh, Y.C. Peng, Z.Y. Lin, PCM based heat sinks of paraffin/ nanoplatelet graphite composite for thermal management of IGBT, Appl. Therm. Eng. 112 (2017) 1129–1136. [2] N.L. Ploch, et al., Effective thermal management in ultraviolet light-emitting diodes with micro-LED Arrays, IEEE Trans. Electron Device 60 (2) (2013) 782–786. [3] X.R. Ding, et al., Thermal and optical investigations of a laser-driven phosphor converter coated on a heat pipe, Appl. Therm. Eng. 148 (2019) 1099–1106. [4] S.W. Zhang, J.L. Chen, Y.L. Sun, J. Li, J. Zeng, W. Yuan, Y. Tang, Experimental study on the thermal performance of a novel ultra-thin aluminum flat heat pipe, Renew. Energy 135 (2019) 1133–1143. [5] M. Janicki, A. Napieralski, Modelling electronic circuit radiation cooling using analytical thermal model, Microelectron. J. 31 (9) (2000) 781–785. [6] S. Belhardj, S. Mimouni, A. Saidane, A. Benzohra, Using microchannels to cool microprocessors: a transmission-line-matrix study, Microelectron. J. 34 (4) (2003) 247–253. [7] Y.X. Wu, et al., Experimental investigation of a PCM-HP heat sink on its thermal performance and anti-thermal-shock capacity for high-power LEDs, Appl. Therm. Eng. 108 (2016) 192–203. [8] J.S. Li, et al., Effect of quantum dot scattering and absorption on the optical performance of white Light-Emitting Diodes, IEEE Trans. Electron. Devices (2018) 1–8. [9] S.M. Sohel Murshed, C.A. Nieto de Castro, A critical review of traditional and emerging techniques and fluids for electronics cooling, Renew. Sustain. Energy Rev. 78 (2017) 821–833. [10] M. Pedram, S. Nazarian, Thermal modeling, analysis, and management in VLSI circuits: principles and methods, Proc. IEEE 94 (8) (2006) 1487–1501. [11] G. Hanreich, J. Nicolics, L. Musiejovsky, High resolution thermal simulation of electronic components, Microelectron. Reliab. 40 (12) (2000) 2069–2076. [12] S.V. Garimella, et al., Thermal challenges in next-generation electronic systems, IEEE Trans. Compon. Packag. Technol 31 (4) (2008) 801–815. [13] C. Yang, D. Liu, F.Y. Zhao, J.F. Tang, Performance analysis and assessment of thermoelectric micro cooler for electronic devices, Energy Convers. Manage 124 (2016) 203–211. [14] D. Pal, M. Severson, Liquid cooled system for aircraft power electronics cooling, Therm. Mech. Phenom. Electron. Syst. IEEE 1 (2017) 800–805. [15] S. Kang, D. Miller, J. Cennamo, Closed loop liquid cooling for high performance computer systems, Am. Soc. Mech. Eng. 2 (2007) 509–515. [16] P.M. Lindh, et al., Direct liquid cooling in low-power electrical machines: proof-ofconcept, IEEE Trans. Energy Convers. 31 (4) (2016) 1-1. [17] R.M. Kumaran, G. Kumaraguruparan, T. Sornakumar, Experimental and numerical studies of header design and inlet/outlet configurations on flow mal-distribution in parallel micro-channels, Appl. Therm. Eng. 58 (1–2) (2013) 205–216. [18] A.O. Ong, A.R. Abramson, N.C. Tien, Electrohydrodynamic microfabricated ionic wind pumps for thermal management applications, J. Heat Transfer 136 (6) (2014) 061703. [19] A. Kamholz, A.E. Kamholz, G. Holman, P. Yager, K.F. Bohringer, A ferrofluidic magnetic micropump, J. Microelectromech. Syst 10 (2) (2001) 215–221.
9
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Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
Research Paper
Experimental investigation on thermal management performance of an integrated heat sink with a piezoelectric micropump
T
Yong Tanga, Mingze Jiaa, Xinrui Dinga, , Zongtao Lia, Zhenping Wana, Qinghong Lina, Ting Fub ⁎
a b
Key Laboratory of Surface Functional Structure Manufacturing of Guangdong High Education Institutes, South China University of Technology, Guangdong, China Key Laboratory of Metallurgical Equipment and Control Technology, Wuhan University of Science and Technology, Wuhan 430081, China
ARTICLE INFO
ABSTRACT
Keywords: Heat sink Piezoelectric micropump Thermal management
High integration density is the trend of future electronics. The accompanying heat dissipation remains a serious problem in electronic thermal management. To overcome the high power density of heat dissipation, an integrated heat sink with a piezoelectric micropump (IHS-PMP) was developed. The driving performance of the IHS-PMP, including the output flow rate and pump pressure, was established to investigate the effects of the driving voltage and driving frequency. Additionally, it was found that the working fluid generated bubbles under a thermal load. The inevitable bubbles significantly affected the flow rate of the IHS-PMP. Finally, the effects of the heating power and driving frequency on the thermal management performance of the IHS-PMP were experimentally investigated and analyzed. The results indicated that with the optimal driving parameters of the IHS-PMP and at a heating power of 80 W, the wall temperature rise was 47.3 °C. This indicates that the IHS-PMP had excellent thermal management performance. The IHS-PMP can be regarded as a high-potential component in loop liquid cooling, which is beneficial to the thermal management of electronic devices.
1. Introduction Thermal management has become a major challenge in the development of electronic components owing to the increased heat flux and reduced component size [1–4]. The reliability of electronic components and optoelectronic materials decreases dramatically as the temperature of the component increases [5–8]. The failure probability of an electronic component increases exponentially when the temperature is greater than 75 °C [9]. According to statistics, more than 50% of electronic component failures are caused by a thermal issue [10]. Additionally, the installation space for thermal management components has become increasingly narrow. The effective thermal management and heat dissipation of electronic devices in a small space has become a bottleneck, restricting the development of electronic devices [11–13]. Loop liquid cooling technology has received extensive attention owing to its stable and efficient long-distance heat transfer, and flexible layout. Pal et al. [14] proposed a liquid cooling system for aircraft power electronic systems and discussed the system-level issues for liquid cooling. Kang et al. [15] introduced the design and properties of an advanced liquid cooling system that can cool single or multiple heat sources in a computer system. The cooling system employs copper cold plates with medium-scale channels to extract heat from the central processing unit and graphics processing unit heat sources, using a ⁎
highly efficient liquid-air heat exchanger with flat copper tubes and common fins to transfer heat to the air via forced convection. A highly reliable and compact pump was used to circulate the fluid in a closed loop. The entire system has a very low fluid permeability and a long service life. However, the loop liquid cooling system contains many components, such as drive pumps, heat sinks, radiators, and pipes. These large components occupy too much cooling space to be used in electronic devices that operate in small spaces. Currently, the driving components of the loop liquid cooling system are predominantly rotor pumps [16,17], which have a loud operating noise, large size, and are not suitable for use in compact loop liquid cooling. Among the numerous micro-driven pumps, the driving abilities of the electrohydrodynamic micropump [18] and the magnetohydrodynamic micropump [19] are weak. The heat-driven micropump [20] is significantly affected by the temperature. The electrostatic actuation micropump [21,22] is difficult to manufacture owing to the use of silicon micromachining and the high precision of the distance between the diaphragm and the electrode. However, the piezoelectric micropump has the advantages of a small temperature influence and a strong driving capability and is very suitable for application in micro-loop liquid cooling systems. Researchers have studied the piezoelectric micropumps used in the loop liquid cooling system [23–28]. Ma et al. [23] replaced the conventional drive
Corresponding author at: South China University of Technology, Guangdong 510641, China. E-mail address: [email protected] (X. Ding).
https://doi.org/10.1016/j.applthermaleng.2019.114053 Received 13 March 2019; Received in revised form 12 June 2019; Accepted 28 June 2019 Available online 29 June 2019 1359-4311/ © 2019 Elsevier Ltd. All rights reserved.
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Nomenclature IHS-PMP PVPC HTC Tin Tout Tci Tw Q ΔV f ξ U ΔV′ P qt qe ql qe′ Ap Rt Rc Rs λc
λs lc
integrated heat sink with a piezoelectric micropump pumping volume per cycle, mL heat transfer coefficient, W m−2 °C−1 inlet fluid temperature, °C outlet fluid temperature, °C thermocouple reading, °C wall temperature, °C theoretical output flow rate without thermal load, mL/min the amount of change in chamber volume due to vibration of the piezoelectric disc in a single cycle, mL frequency of sinusoidal AC power supply, Hz piezoelectric drive efficiency coefficient driving voltage, V pumping volume per cycle, mL pump pressure, kPa total heating power, W effective heating power, W loss of heating power, W effective heat flux supplied to test section, kW/m2 The ratio of the loss of heating power to the total heating power area of the heat exchange contact surface, mm2 total thermal conduction resistance, °C/m thermal conduction resistance of pure copper, °C/m thermal conduction resistance of solder, °C/m thermal conductivity of pure copper, W m−1 °C−1
ls ΔT Ta R’ h Q’ c ρ
thermal conductivity of solder, W m−1 °C−1 distance between the upper surface of the copper block and the thermocouple, mm thickness of the solder layer, mm wall temperature rise, °C average temperature of the inlet and outlet liquid temperature, °C thermal resistance of the IHS-PMP, °C/W heat transfer coefficient, W m−2 °C−1 flow rate of IHS-PMP in the thermal management test, mL/min specific heat capacity of deionized water, J kg−1 °C−1 density of deionized water, g/cm3
Subscript in out ci w t e l p c s a
pump with a piezoelectric micropump, which was connected to a heat sink to form a new liquid cooling system. It satisfies the system miniaturization requirements and achieves good thermal management performance. Researchers have improved the structure of piezoelectric micropumps. Ma et al. [24] improved the full-side support mode of the piezoelectric vibrator and proposed a single-sided support cantilever assembly method, which can increase the effect of the piezoelectric vibrator on the volume of the pump chamber. The specific effect is to increase the volume change. Shabanian et al. [25] improved the
inlet outlet thermocouple location in stream-wise direction wall total effective loss platform of on the upper surface of the copper block copper solder average
structure of the piezoelectric vibrator, proposing a ring-shaped structure, which improved the displacement and the output performance of the piezoelectric micropump. Ren et al. [26] improved the check valve structure and proposed a long diaphragm-shaped check valve, which can significantly improve the operating frequency of the piezoelectric pump. Additionally, other researchers have proposed miniaturizing the heat sink in the liquid cooling of the loop to further reduce the size of the radiator [27,28]. The aforementioned studies were aimed at miniaturizing the entire
Fig. 1. Diagram of IHS-PMP ((a) is the structure diagram of IHS-PMP. (b) is the structure diagram of the heat exchange chamber). 2
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coolant (deionized water) to flow uniaxially through the heat exchange chamber. The drive unit was composed of a piezoelectric disc and a double O-ring. The piezoelectric disc (P-58, peak-to-peak value of the maximum allowable driving voltage was 300 V) was composed of a piezoelectric ceramic and a brass foil bonded together. The piezoelectric ceramic had a diameter of 36 mm and a thickness of 0.38 mm. The brass foil had a diameter of 41 mm and a thickness of 0.18 mm. One electrode of the piezoelectric disc was soldered to the piezoelectric ceramic, and the other was soldered to the brass foil. In this study, it was necessary that the piezoelectric disc operated in a heated state, and the sealing property was more demanding. Therefore, a double O-ring sealing element structure was designed. The double O-ring was similar to the overlap of two O-rings, except that there was a 0.2-mm-thick connection between the two O-rings, improving the sealing performance. The piezoelectric disc was clamped by a double O-ring with a diameter of 42 mm to form the drive unit. A pressing plate with a thickness of 3 mm was used to compress the double O-ring of the drive unit. The stainless-steel gasket was used to limit the amount of depression of the pressing plate.
heat dissipation system by reducing the size of various components in the loop liquid cooling. However, complex system components in the loop liquid cooling system lead to increased connection points of various components, which may lead to leakage of the working fluid, resulting in system failure or even damage to the microelectronic devices. Therefore, the development of new heat dissipation components and the improvement of the integration of the loop liquid cooling system to satisfy the high heat flux thermal management requirements in the confined spaces of electronic devices are urgently needed. Integration technology reduces the number of components and realizes the multifunctional application of components [29–31]. The integration of a piezoelectric micropump and a micro-heat sink is a feasible solution for effective thermal management. At present, the radiator is integrated with the vapor chamber [32–34] and heat pipe [35,36], which can make the radiator structure compact, reduce the thermal resistance, and improve the thermal management performance. However, there has been little research on the integration of an efficient piezoelectric pump with a micro heat sink. In this work, a novel integrated heat sink with a piezoelectric micropump (IHS-PMP) is proposed to satisfy the high heat flux thermal management requirements in the confined spaces of electronic devices. The IHS-PMP driving performance without a thermal load and the thermal management performance were experimentally evaluated.
3. Experimental setup and procedure In this work, two measurements were performed to synthetically investigate the performance of the IHS-PMP, the driving performance without a thermal load (including the output flow rate measurement and pump pressure measurement), and the thermal management performance (including the temperature rising and thermal resistance measurements). The device schematic is shown in Fig. 2. The power drive part include an arbitrary waveform generator (AFG3251, maximum output voltage amplitude of 5 V) and a power amplifier (ATA2210, maximum magnification of 100). Deionized water was utilized as the working liquid. The deionized water was in a constant temperature water tank with an automatic temperature regulation system. The temperature of the constant temperature water tank was set as 25 °C. The driving performance test platform is shown in Fig. 2(a). As the
2. Design and fabrication of IHS-PMP A diagram of the IHS-PMP is shown in Fig. 1. The IHS-PMP comprised four parts: a heat exchange chamber, a drive unit, a stainless steel gasket, and a pressing plate. The copper heat exchange chamber was fabricated by milling a cylinder with a depth of 4.8 mm and a diameter of 33 mm. The design principle of the IHS-PMP heat exchange chamber was to use a part of the piezoelectric pump as a heat sink; that is, the heat sink and the piezoelectric pump shared a chamber for integration purposes. Thus, the IHS-PMP had the functions of pumping liquid and heat transfer. Two check valves were utilized to cause the
Fig. 2. Diagram of experimental setup ((a) is the flow rate and pump pressure test device diagram of IHS-PMP without thermal load. (b) is the thermal management performance test device diagram of IHS-PMP). 3
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series flow meter affected the pump pressure of the system, it also further affect the flow rate. The flow rate test section comprised a beaker, an electronic balance, and a timer. The flow rate of the IHSPMP without a thermal load was measured by averaging the pumped water mass within 1 min. Two lifting platforms were used to adjust the water tank height and output piping to ensure that the water tank, output piping, and IHS-PMP had the same height. The pressure test part comprised a differential pressure transmitter, a direct-current (DC) stabilized power source, a pressure relief valve, and a temperature acquisition card. The outlet of the IHS-PMP was connected to the differential pressure transmitter (PCM300, range of 0–20 kPa) and the pressure relief valve via a three-way valve; the pressure relief valve was used to vent the pressure in the tube after the test was completed. Similarly, the tank, IHS-PMP, and pressure transmitter were at the same height. The DC stabilized power source supplied power to the differential pressure transmitter and the temperature acquisition card. Once the pressure was steady, it was recorded by the pressure transmitter with a sampling rate of 1 Hz for a period of 1 min. A diagram of the thermal management performance test part is shown in Fig. 2(b). The constant-temperature (25 °C) deionized water in the water tank was circulated by the HIS-PMP. A cross-sectional view of the thermal management performance test device is shown in Fig. 3. Nine heaters were embedded inside a copper cylinder with a diameter of 33 mm. The test sample was soldered on top of the copper block using a thin layer of solder (Pb-Sn-Ag-Sb) to minimize the contact thermal resistance. The nine heaters were powered by a variac and a wattmeter showed the input power. The heating area was embedded into a thermal insulation material (thermal conductivity of ≤0.27 W/ m °C) to reduce the heat loss. In the data acquisition part, seven K-type thermocouples were employed to measure the temperature, with two at the inlet/outlet of the IHS-PMP (Tin, Tout), and three (T1, T2, T3,) were distributed 5 mm below the test surface with intervals of 12 mm. In the vertical direction, two K-type thermocouples (T4, T5) were placed below T2 with an interval of 6 mm. A HIOKI data acquisition system (LR841030 wireless logging station and LR8510 wireless voltage/temp unit) was employed to collect and display the temperature using the seven K-type thermocouples. The wall temperature (Tw) was then calculated from the Fourier Law.
4. Data reduction 4.1. Theoretical output flow rate and output pressure of IHS-PMP The design of the IHS-PMP is based on the principle of a piezoelectric micropump and its output flow rate can be calculated as follows: (1)
Q = 60 Vf
where ΔV is the amount of change in the chamber volume due to the vibration of the piezoelectric disc in a single cycle, f is the frequency of the sinusoidal alternating-current power supply, and ξ is the piezoelectric drive efficiency coefficient. Stemme et al. [37,38] proposed a method for calculating ΔV. The relationship between ΔV and U is:
V
(2)
U
where U is the driving voltage applied to the piezoelectric disc. To facilitate the study of the relationship between the driving frequency and the output flow rate, the pumping volume per cycle (PVPC) of the IHS-MPM is defined as the amount of liquid volume pumped by the IHS-MPM when the piezoelectric disc vibrates for one cycle and is denoted as ΔV′:
V' = V =
Q 60f
(3)
The working fluid used in this experiment was deionized water, which can be considered as an incompressible fluid. According to the assumption that the piezoelectric drive pump chamber is a closed space [37,38], the relationship between the pump pressure P and U is:
P
U
(4)
4.2. Wall temperature of IHS-PMP is:
The total heating power qt in the thermal management experiment
qt = qe + ql
(5)
where qe is the effective heating power applied to the heated surface
Fig. 3. Details of the thermal management performance test device (Rc is the thermal conduction resistance of the copper block, Rs is the thermal conduction resistance of the solder layer, Rt is the total thermal conduction resistance, R' is the thermal resistance of the IHS-PMP). 4
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and ql is the loss of heating power. An insulating cotton with a thickness of 20 mm was wrapped around the outside of the actual test platform. The one-dimensional Fourier Law was employed to calculate the actual heat flux densityqe' according to T2, T4, and T5:
qe' =
c
ql = qt
dT dx
(6)
qe' Ap
(7)
h=
ql qt
Q' =
qe
(9)
qe (10)
Ap
Because the direction of the heat flux flow was straight upward, the wall temperature was defined as the surface temperature of the IHSPMP, which was the temperature of the upper surface of the solder. Generally, in the IHS-PMP chamber, the temperature of the fluid increased linearly along the flow direction. According to the one-dimensional heat conduction equation, we define the temperature at the midpoint of the IHS-PMP as the wall temperature Tw:
Tw = T2
The driving frequency and driving voltage significantly affected the output flow rate and pump pressure of the piezoelectric micropump. The output flow rate of the IHS-PMP had a large influence on the thermal management performance. Therefore, before the thermal management performance test of the IHS-PMP, it was necessary to study its driving performance without a thermal load, focusing on the effects of the driving voltage and driving frequency on the output flow rate of the IHS-PMP. The thermal management performance test of the IHS-PMP was based on the results of the driving performance without a thermal load.
where Rt represents the total thermal conduction resistance along the heat conduction direction, which can be calculated as follows: (12)
Rt = R c + Rs
where Rc represents the thermal conduction resistance of the copper block, and Rs represents the thermal conduction resistance of the solder layer:
Rc =
lc , Rs = c Ap
ls s Ap
5.1. Driving performance of IHS-PMP without thermal load The driving voltages of 200 and 300 V were selected to study the influence of the driving frequency on the driving performance of the IHS-PMP. The check valve response frequency is generally within 100 Hz; thus, the driving frequency was optimized in the frequency range of 0–100 Hz. The driving frequency parameters were set in the range of 0–100 Hz, with intervals of 5 Hz. Fig. 4 shows the output flow rate of the IHS-PMP. The maximum output flow rates for 200 and 300 V (140 and 210 mL/min, respectively) occurred at the frequency of 15 Hz. In the driving frequency range of 5–10 Hz, the piezoelectric disc vibrated at a lower frequency. As the driving frequency increased, the check valve gradually opened and the output flow rate increased. This is illustrated in Fig. 5; as the frequency increased, the PVPC increased. The PVPC reached its maximum value, and the check valve was fully open at 10 Hz. In the driving frequency range of 10–15 Hz, because the motion of the check valve lagged behind the vibration of the piezoelectric disc, the PVPC decreased. However, as the vibration frequency of the piezoelectric disc increased, the output flow rate of the IHS-PMP slowly increased. The maximum output flow rate was reached at 15 Hz. At 15–25 Hz, the PVPC decreased sharply owing to the lagging motion of the check valve behind the vibration of the piezoelectric disc. Even though the increasing driving frequency could increase the total number of pumping cycles, it was insufficient to compensate for the reduction in the PVPC. The output flow rate was gradually reduced at 15–25 Hz. When the driving frequency increased to 25–70 Hz, the second maximum of the output flow rate occurred in the driving frequency
(13)
Here, lc = 5 mm represents the distance between the upper surface of the copper block and the thermocouple (T1, T2, or T3), λc = 397 W/ (m °C) represents the thermal conductivity of pure copper, ls = 0.2 mm represents the thickness of the solder layer, and λs = 50 W/(m·°C) represents the thermal conductivity of the solder. The wall temperature rise is defined as:
T = Tw
Tin
(14)
where Tin represents the temperature of the thermocouple at the entrance. 4.3. Heat transfer coefficient (HTC) of IHS-PMP The average of the inlet and outlet liquid temperatures was taken as the fluid temperature:
Ta =
Tin + Tout 2
(15)
where Tout represents the temperature of the thermocouple at the outlet. The thermal resistance of the IHS-PMP can be calculated as follows:
R' =
(Tw
Ta ) qe
(18)
5. Results and discussion
(11)
qe Rt
Tin)
The dimensional errors due to the limitation of the milling accuracy and the solder layer thickness were estimated to be ± 0.01 mm. The accuracy of the temperature measurement by the precise K-type thermocouples was ± 0.2 K; The uncertainty of the power input was ± 1 W; The uncertainty of the pressure transmitter was ± 12 Pa; The uncertainty of the electronic balance was ± 0.01 g; Using a standard error analysis, the calculated uncertainties for the heat flux, the HTC, and the flow rate in the thermal management performance test were within 1.7%, 2.6%, and 1.6%, respectively [39].
The effective heat flux is:
qe' =
qe c (Tout
4.4. Uncertainties
in this experiment was determined to
qt
(17)
Ta
where c = 4186 J/(kg °C) represents the specific heat capacity of deionized water, and ρ = 1.0 g/cm3 represents the density of deionized water.
(8)
After testing, the be < 5%.Thus, ql ≪ 8qt.
Tw
To avoid the impact of the serial flowmeter on the IHS-PMP pressure and flow rate, in the thermal management test, the flow rate of the IHSPMP was calculated as follows:
where Ap is the area of the heat exchange contact surface (33 mm × 33 mm). The ratio of the loss of heating power to the total heating power is:
=
qe'
(16)
The HTC is calculated as follows: 5
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15, 20, 40, 45, and 50 Hz. According to the foregoing discussion, the two frequencies ƒ1 (15 Hz) and ƒ2 (45 Hz) exhibited reasonable output flow rates. Therefore, ƒ1 (15 Hz) and ƒ2 (45 Hz) were selected to investigate the influence of the driving voltage. The operating driving voltage ranged from 100 to 300 V, with intervals of 50 V. The IHS-PMP output flow rate at both driving frequencies increased linearly with the increasing driving voltage, as shown in Fig. 7. As the driving voltage increased, the deformation of the piezoelectric disc and the cavity change volume ΔV increased linearly, resulting in a linear increase in the output flow rate of the IHS-PMP. This is theoretically consistent with the relationship between the driving voltage and the output flow rate [38]. However, the driving voltage cannot be increased indefinitely. The maximum operating voltage of the piezoelectric disc selected in this study was 300 V. The IHS-PMP achieved the maximum output flow rate at 300 V. Additionally, the IHS-PMP pump pressure at both driving frequencies increased linearly with the increasing driving voltage, as shown in Fig. 7. This is theoretically consistent with the relationship between the driving voltage and the pump pressure [38]. Similarly, owing to the limitation of the operating voltage of the piezoelectric disc, the IHSPMP achieved the maximum pump pressure at a driving voltage of 300 V in this study.
Fig. 4. Flow rate and driving frequency relationship at different driving voltages.
5.2. Thermal management performance of IHS-PMP The thermal management performance of the IHS-PMP was tested under different driving frequencies. Six driving frequencies were selected with relatively high output flow rates, as previously mentioned. The driving voltage of 300 V with the maximum output flow rate was selected. The output heating power was increased from 0 W with intervals of 10 W. When the temperature test points T1–T5 fluctuated less than 0.5 °C within 1 min, the IHS-PMP was considered to be in a steady state. The upper limit working temperature was 70 °C. When the temperature test points T1-T5 of the IHS-PMP were > 70 °C, the IHS-PMP was considered to have reached the working limit, and the experiment was terminated. When the temperature of the IHS-PMP reached the upper limit working temperature, the heating power was defined as the upper limit heating power. If the IHS-PMP temperature fluctuated greatly and could not be stabilized within 15 min, the IHS-PMP was considered to have reached the stable limit working temperature and stopped heating. When the temperature of the IHS-PMP reached the limit working temperature, the heating power was defined as the stable limit working power. Fig. 8 shows the wall temperature rise of the IHS-PMP at the driving frequencies of 10, 15, 20, 40, 45, and 50 Hz. The wall temperature rise
Fig. 5. PVPC and driving frequency relationship at different driving voltages.
range of 45 Hz. When the driving frequency was 45 Hz, for the driving voltages of 200 and 300 V, the output flow rates of the IHS-PMP were 125 and 180 mL/min, respectively. In the driving-frequency range of 25–45 Hz, as the driving frequency increased, the phase lag of the check valve with respect to the piezoelectric disc exceeded one cycle; thus, the check valve was matched with the next cycle of the piezoelectric disc, and the output flow rate started to increase again. According to a previous study on the existing piezoelectric micropump [40], in the driving-frequency range of 45–70 Hz, the check valve cannot be fully opened, and the output flow rate decreases. Moreover, as the driving frequency increases, the minimum distance between the valve and the valve seat increases, and the phenomenon of incomplete closing is more pronounced, which results in a decrease in the output flow rate. Therefore, in the driving-frequency range of 25–70 Hz, a peak of the flow rate occurred when the driving frequency was 45 Hz. In the frequency range of 70–100 Hz, although the flow rate increased slightly at 80 Hz, the output flow rate was relatively low, the vibration frequency of the piezoelectric disc was too high, and the selected check valve did not effectively match the work. Therefore, this driving frequency range cannot be used for IHS-PMP operation. Fig. 6 shows the output pump pressure of the IHS-PMP in the driving frequency range of 5–100 Hz. The output pump pressure remained approximately constant at a high level. This indicates that the driving frequency had little effect on the IHS-PMP pump pressure. Therefore, the appropriate driving frequencies for the IHS-PMP samples were 10,
Fig. 6. Pump pressure and driving frequency relationship at different driving voltages. 6
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decreased as the heat power increased. At the driving frequency of 10 Hz, the flow rate decreased from 167.01 to 61.48 mL/min as the heating power increased from 10 to 80 W. Even the flow rate at 10 W was lower than the flow rate without heating. The decrease in the flow rate was caused by the generation of bubbles in the system, even though the water in the circulating water tank was degassed. Traditional thermal management performance tests of a heat sink are only executed at a certain flow rate. When the flow rate is chosen, it will not change with the increasing heat power. The actuating device in the loop system increases its power to maintain a stable flow rate when the heat power is increased. In this study, the thermal management performance tests of the IHS-PMP were performed under a constant actuating power. Even at the maximum heating power, the water temperature was < 73 °C; thus, the bubbles were not generated by the phase change. The thermal management performance test in this study was conducted in an open environment, which simulated the scene of the heat sink in the actual work. Before the test, the deionized water in the constanttemperature water tank was boiled sufficiently to remove non-condensable gases. However, as the test was conducted in an open environment, the environment of the constant temperature water tank could not be a vacuum; thus, some non-condensable gases were still dissolved in the water. When deionized water passed through the heat exchange chamber, the temperature difference between the deionized water and the wall changed the solubility of the gas in the water. Therefore, the gas precipitated out of the fluid, increasing the fluid resistance. This phenomenon occurs in a loop liquid cooling system operating in a non-vacuum environment [45,46]. The bubbles had a certain incompressibility which reduced the volume of the micropump chamber when the piezoelectric disc vibrated. This reduced the output flow rate of the IHS-PMP. Additionally, as the heating power increased, the temperature difference between the deionized water and the wall increased, which generated more bubbles, reducing the flow rate. Thus, the IHS-PMP followed the same principle (of flow-rate reduction), as the traditional piezoelectric micropump chamber [24]. This phenomenon is not caused by integration; it also occurs if the driving pump and the heat sink work in series. In the thermal management performance test of a traditional cooling system, the power of the driving pump is always increased to maintain the system output flow rate stability, but in this study, the driving part was integrated with the heat transfer part; thus, the driving parameters were not changed in the thermal management performance test. The IHS-PMP has a different ability to discharge air bubbles from the heat exchange chamber at different driving frequencies. When the driving frequency was 10 Hz, the pumping amount and output flow rate were maximized, and the bubbles were effectively discharged in the heat exchange chamber. Additionally, it was observed that the bubbles
Fig. 7. Effect of the driving voltage on the flow rate and pump pressure of the IHS-PMP.
Fig. 8. ΔT and heating power relationship at different driving frequencies.
was smallest in the driving frequency range of 10 Hz. As the driving frequency increased, the wall temperature rise also increased, but the wall temperature rise was greatest at 20 Hz. For instance, when the heating power was 40 W and the driving frequency was 10, 15, 20, 40, 45, and 50 Hz, the wall temperature rise of the IHS-PMP was 14.75, 16.73, 31.04, 20.66, 21.65, and 27.11 °C, respectively. Additionally, the ultimate power at different driving frequencies is shown in Fig. 9. The ultimate power of the IHS-PMP at 10, 15, 20, 40, 45, and 50 Hz was 80, 60, 40, 50, 50, and 40 W, respectively. The ultimate power of the IHSPMP at 10 Hz was limited by the upper limit working temperature, and at other driving frequencies, it was limited by the heat transfer stabilization before the test temperature reached the stable limit working temperature. The most suitable driving frequency for the IHS-PMP was 10 Hz. Compared with the thermal management method using aircooling [41] and phase-change heat transfer [32,42], under the same heating conditions, the thermal management performance of the IHSPMP in the loop liquid cooling was excellent at the driving frequency of 10 Hz. When the driving frequency was 10 Hz, compared with other thermal management methods using loop liquid cooling [43,44], the thermal management performance of the IHS-PMP was excellent at the same volume of the heat sink. The differences in the thermal management performances of the IHS-PMP were caused by the different driving frequencies, because at different driving frequencies, as the thermal power increased, the flow rate changed differently. Fig. 10 shows the variation of the IHS-PMP flow rate with the increasing thermal power at different driving frequencies. The flow rate
Fig. 9. Ultimate power at different driving frequencies. 7
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Fig. 12. HTC and heating power relationship at different driving frequencies.
Fig. 10. Flow rate and heating power relationship at different driving frequencies.
stably discharged. This is because at 10 Hz, the IHS-PMP had the lowest vibration frequency, which improved the stability of degassing. Although the driving frequencies of 15, 20, 40, 45, and 50 Hz yielded a considerable flow rate without a thermal load, the ability of the IHSPMP to discharge air bubbles was poor owing to the small PVPC. At the driving frequencies of 15 and 40 Hz, the bubbles oscillated back and forth at the outlet pipe. At the driving frequencies of 20, 45, and 50 Hz, as the heating power increased, the bubbles became larger and increased in number but did not move. The bubble movements in the chamber were considered to be similar. Bubbles that could not be discharged from IHS-PMP accumulated in the heat-exchange chamber. Therefore, at the same power, the amount of residual bubbles in the chamber was minimized when the IHS-PMP was at 10 Hz, the output flow rate was maximized, and the wall temperature rise was minimized. Because of the bubble generation and the decrease in the output flow rate, the thermal resistance of the IHS-PMP increased with the increasing heat power, as shown in Fig. 11. Compared with other driving frequencies, the thermal resistance was the lowest at 10 Hz. The IHSPMP was in the most efficient heat transfer state at 10 Hz. However, the overall thermal resistance of IHS-PMP is still relatively high, because inside the IHS-PMP, the water flow is considered to be laminar, and the flow velocity at the bottom of the heat exchange chamber is small. This part of the water directly exchanges heat with the heat exchange chamber, which will result in a relatively large thermal resistance of the IHS-PMP.
The HTC of the IHS-PMP is shown in Fig. 12. The IHS-PMP had the largest HTC at the driving frequency of 10 Hz, because the maximum output flow rate was 10 Hz with the minimal thermal resistance. Moreover, as the heating power increased, the HTC gradually decreased, because the flow rate gradually decreased. Additionally, 10 Hz yielded the optimum flow-rate stability, which is advantageous for heat exchange. Therefore, 10 Hz is the optimal driving frequency for heat exchange. At other driving frequencies, the output flow rate of the IHSPMP was low, and the liquid flow was unstable. This explains why the temperature of the IHS-PMP at other driving frequencies could not reach the upper limit working temperature. 6. Conclusion An IHS-PMP was proposed to address the thermal management demand of electronic devices in narrow spaces. The IHS-PMP solved the problems of large operating noise and a large rotor pump in loop liquid cooling. [16,17] Additionally, the problem of the leakage of working fluid caused by the piezoelectric micropump and the heat sink in series was solved by integration. [23–28]. The compactness of the loop liquid cooling system was improved via integration of a pump and a heat sink. First, the driving performance of the IHS-PMP without a thermal load was investigated. Then, the thermal management performance of the IHS-PMP was evaluated. The following conclusions were drawn.
• The output flow rate of the IHS-PMP without a thermal load was • •
significantly affected by the driving frequency owing to the influence of the matching of the piezoelectric disc and the check valve. The output pressure of the IHS-PMP without a thermal load was essentially unaffected by the driving frequency. The output flow rate and output pressure of the IHS-PMP without a thermal load were proportional to the driving voltage. The thermal management performance of the IHS-PMP was affected by not only the output flow rate but also the generation and discharge of bubbles in the heat exchange chamber. A high output flow rate and effective bubble discharge contributed to the thermal management performance of the IHS-PMP. The driving frequency of 10 Hz and driving voltage of 300 V were the best thermal management driving parameters for the IHS-PMP. Under these conditions, when the heating power of the IHS-PMP was 80 W, the wall temperature rise was 47.3 °C, which is suitable for the thermal management of electronic devices in narrow spaces.
Although the working mechanism and thermal management performance of IHS-PMPs have been studied, it is necessary to study the
Fig. 11. R and heating power relationship at different driving frequencies. 8
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structure of the heat exchange chamber and the bubble dynamics in the liquid to improve the heat management effect.
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Acknowledgements This work was supported by the National Natural Science Foundation of China (No. 51805173, No. 51735004), Science and Technology Program of Guangzhou, China (No. 201904010252), Fundamental Research Funds for the Central Universities (No. 2018MS44), The Open Fund of the Key Laboratory for Metallurgical Equipment and Control of Ministry of Education in Wuhan University of Science and Technology (No. 2018B01), Science and Technology Planning Project of Guangdong Province, China (No. 2015B020238004). Appendix A. Supplementary material Supplementary data to this article can be found online at https:// doi.org/10.1016/j.applthermaleng.2019.114053. References [1] T.C. Chang, S. Lee, Y.K. Fuh, Y.C. Peng, Z.Y. Lin, PCM based heat sinks of paraffin/ nanoplatelet graphite composite for thermal management of IGBT, Appl. Therm. Eng. 112 (2017) 1129–1136. [2] N.L. Ploch, et al., Effective thermal management in ultraviolet light-emitting diodes with micro-LED Arrays, IEEE Trans. Electron Device 60 (2) (2013) 782–786. [3] X.R. Ding, et al., Thermal and optical investigations of a laser-driven phosphor converter coated on a heat pipe, Appl. Therm. Eng. 148 (2019) 1099–1106. [4] S.W. Zhang, J.L. Chen, Y.L. Sun, J. Li, J. Zeng, W. Yuan, Y. Tang, Experimental study on the thermal performance of a novel ultra-thin aluminum flat heat pipe, Renew. Energy 135 (2019) 1133–1143. [5] M. Janicki, A. Napieralski, Modelling electronic circuit radiation cooling using analytical thermal model, Microelectron. J. 31 (9) (2000) 781–785. [6] S. Belhardj, S. Mimouni, A. Saidane, A. Benzohra, Using microchannels to cool microprocessors: a transmission-line-matrix study, Microelectron. J. 34 (4) (2003) 247–253. [7] Y.X. Wu, et al., Experimental investigation of a PCM-HP heat sink on its thermal performance and anti-thermal-shock capacity for high-power LEDs, Appl. Therm. Eng. 108 (2016) 192–203. [8] J.S. Li, et al., Effect of quantum dot scattering and absorption on the optical performance of white Light-Emitting Diodes, IEEE Trans. Electron. Devices (2018) 1–8. [9] S.M. Sohel Murshed, C.A. Nieto de Castro, A critical review of traditional and emerging techniques and fluids for electronics cooling, Renew. Sustain. Energy Rev. 78 (2017) 821–833. [10] M. Pedram, S. Nazarian, Thermal modeling, analysis, and management in VLSI circuits: principles and methods, Proc. IEEE 94 (8) (2006) 1487–1501. [11] G. Hanreich, J. Nicolics, L. Musiejovsky, High resolution thermal simulation of electronic components, Microelectron. Reliab. 40 (12) (2000) 2069–2076. [12] S.V. Garimella, et al., Thermal challenges in next-generation electronic systems, IEEE Trans. Compon. Packag. Technol 31 (4) (2008) 801–815. [13] C. Yang, D. Liu, F.Y. Zhao, J.F. Tang, Performance analysis and assessment of thermoelectric micro cooler for electronic devices, Energy Convers. Manage 124 (2016) 203–211. [14] D. Pal, M. Severson, Liquid cooled system for aircraft power electronics cooling, Therm. Mech. Phenom. Electron. Syst. IEEE 1 (2017) 800–805. [15] S. Kang, D. Miller, J. Cennamo, Closed loop liquid cooling for high performance computer systems, Am. Soc. Mech. Eng. 2 (2007) 509–515. [16] P.M. Lindh, et al., Direct liquid cooling in low-power electrical machines: proof-ofconcept, IEEE Trans. Energy Convers. 31 (4) (2016) 1-1. [17] R.M. Kumaran, G. Kumaraguruparan, T. Sornakumar, Experimental and numerical studies of header design and inlet/outlet configurations on flow mal-distribution in parallel micro-channels, Appl. Therm. Eng. 58 (1–2) (2013) 205–216. [18] A.O. Ong, A.R. Abramson, N.C. Tien, Electrohydrodynamic microfabricated ionic wind pumps for thermal management applications, J. Heat Transfer 136 (6) (2014) 061703. [19] A. Kamholz, A.E. Kamholz, G. Holman, P. Yager, K.F. Bohringer, A ferrofluidic magnetic micropump, J. Microelectromech. Syst 10 (2) (2001) 215–221.
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