Cryogenic spray quenching of simulated propellant tank wall using coating and flow pulsing in microgravity.

In-space cryogenic propulsion will play a vital role in NASA's return to the Moon mission and future mission to Mars. The enabling of in-space cryogenic engines and cryogenic fuel depots for these future manned and robotic space exploration missions begins with the technology development of advanced cryogenic thermal-fluid management systems for the propellant transfer lines and storage system. Before single-phase liquid can flow to the engine or spacecraft receiver tank, the connecting transfer line and storage tank must first be chilled down to cryogenic temperatures. The most direct and simplest method to quench the line and the tank is to use the cold propellant itself that results in the requirement of minimizing propellant consumption during chilldown. In view of the needs stated above, a highly efficient thermal-fluid management technology must be developed to consume the minimum amount of cryogen during chilldown of a transfer line and a storage tank. In this paper, we suggest the use of the cryogenic spray for storage tank chilldown. We have successfully demonstrated its feasibility and high efficiency in a simulated space microgravity condition. In order to maximize the storage tank chilldown efficiency for the least amount of cryogen consumption, the technology adopted included cryogenic spray cooling, Teflon thin-film coating of the simulated tank surface, and spray flow pulsing. The completed flight experiments successfully demonstrated that spray cooling is the most efficient cooling method for the tank chilldown in microgravity. In microgravity, Teflon coating alone can improve the efficiency up to 72% and the efficiency can be improved up to 59% by flow pulsing alone. However, Teflon coating together with flow pulsing was found to substantially enhance the chilldown efficiency in microgravity for up to 113%.

Introduction

In NASA’s return to the Moon mission and future mission to Mars, a highly efficient cryogenic thermal-fluid management technology is among the indispensable requirements for successful lunar and mars space missions. The planned propellant fuel depot deployed in the Lower-Earth-Orbit (LEO) for future deep-space missions, and the human-carrying spacecraft flying lunar and mars missions are designed to utilize liquid cryogenic fuels and oxydizers[1-4]. For the human mars surface mission, one of the enabling technologies is the efficient transfer of propellant from the fuel depot to the spacecraft propellant storage tank[1]. The actual tank-to-tank propellant transfer, however, has yet to take place, mainly due to the lack of cryogenic quenching data in reduced microgravity[4] for designing the transfer system. As the existing technology on cryogenic chilldown can only offer relatively very low efficiencies[5] and it has not been developed under the microgravity conditions, a new technology with much higher efficiencies and verified under microgravity conditions is therefore needed for future space missions.

In order to maximize the storage tank chilldown efficiency, the technology proposed includes cryogenic spray cooling, Teflon thin-film coating of the tank inner surface, and spray flow pulsing. The completed flight experiments successfully demonstrated that cryogenic spray cooling is the most efficient cooling method for the tank wall chilldown in microgravity. Teflon coating together with flow pulsing was found to substantially enhance the chilldown efficiency in microgravity. The feasibilities of charge-hold-vent for tank chilldown and no-vent-fill for tank filling in microgravity were also successfully demonstrated.

According to publications by the members of the Space Cryogenic Thermal Management Group at NASA Glenn Research Center, Doherty et al.[6] and Myer et al.[7] provided the main areas of research and development for space cryogenic thermal management. The tank-to-tank transfer of propellants in space is composed of transfer line chilldown, receiver tank chilldown, and no-vent fill of the receiver tank. Among all three areas, receiver tank chilldown is considered the most important area as the amount of cryogen consumed is the largest. In this paper, we report an advance in microgravity tank wall chilldown heat transfer using a thin-film coating and spray cooling.

The chilldown of the receiver tank wall by spray cooling using liquid cryogen is a liquid-to-vapor phase change quenching process that is characterized by the so-called “boiling curve” as shown in Fig. 1. This curve[8] represents the tank wall surface heat flux,, plotted against the wall surface degree of superheating, , where is the surface temperature and is the saturation temperature corresponding to the boiling fluid bulk pressure. A quenching process follows the route D→C→B→A. Therefore, during chilldown the heat transfer on the tank wall surface always experiences film boiling first due to a very hot tank wall surface. Because the heat fluxes in film boiling are relatively quite low, film boiling regime always occupies the major portion of the total quenching time. Accordingly, the thermal energy efficiency in the traditional quenching process is extremely low. According to Shaeffer et al.[5], the average quenching efficiency is about 8% that provides a strong incentive to find more efficient methods for the space storage tank chilldown process.

Fig. 1

A typical boiling curve.

This cure illustrates different boiling regimes and corresponding flow patterns.

Progressive advances in high power density electronics and high-performance energy systems have precipitated the need for innovative thermal management technologies to ensure reliable performance and reduce the payload of thermal management systems. Such systems include high current density propulsion systems, high power electronics for energy conversion, high power optical sensors, as well as high power microelectronics packaged within environmental enclosures. In order to manage the progressively increasing heat flux requirements for thermal management systems, a spray cooling system has been proposed and under constant development for the past sixty years. Liang and Mudawar[9] indicated that spray cooling possesses several advantages: high flux heat dissipation, low and fairly uniform surface temperature, and ability to cool relatively large surface areas using a single nozzle.

In very recent papers, Liang and Mudawar[9,10] provided a highly comprehensive, thorough, and complete review of the spray cooling research up to 2017. Almost all of the published papers were using water and refrigerants and we found only three papers on the study of terrestrial cryogenic spray cooling. Sehmbey et al.[11] performed a liquid nitrogen spray cooling experiment to gather heat transfer characteristics to facilitate the operation of power electronics at very low temperatures. Four different nozzles at various pressures were used to study the variation in spray cooling heat transfer at liquid nitrogen temperature. The effect of nozzle and flow rate on the critical heat flux (CHF) and overall heat transfer characteristics were presented. Cooling heat fluxes close to 1.7 × 106 W/m2 were realized at temperatures below 100 K. The mass flow rate range was from 6.1 × 104 to 3.2 × 105 kg/h m2. They demonstrated that a high heat flux (over 1.0 × 106 W/m2) cooling technique, such as spray cooling, will have to be used to realize all the advantages of low-temperature operation. Following their experimental study, Sehmbey et al.[12] further provided empirical correlations for liquid nitrogen spray cooling. They offered a general semiempirical correlation (based on macrolayer dryout model) for spray cooling CHF for different liquids and spray conditions. An empirical correlation for heat flux was also presented. They also pointed out the importance of surface roughness for spray cooling with liquid nitrogen. It was discovered that the rougher surfaces have significantly higher heat transfer rates and similar CHFs occurring at lower temperatures.

Somasundara and Tay[13] investigated the intermittent liquid nitrogen spray cooling for applications, which require higher heater operating temperatures (−180 to 20 °C). This intermittent spray cooling process can be adjusted using the mass flow rate, pulsing frequency, and duty cycle (percentage of open time in one cycle) to match the required cooling rate on the target. The intermittent spray experiments were conducted for various ranges of surface temperatures.

Kato et al.[14] studied the gravity effects on liquid spray cooling using a nickel-plated copper block in terrestrial and variable gravity conditions onboard parabolic flight. The copper block was first heated by seven cartridge heaters to a prescribed temperature and then cooled down by spraying water or CFC-113 onto the nickel-plated surface which is only 19 mm in diameter. They observed that the heat transfer in the low heat flux regime below the CHF was enhanced by the reduction in gravity for both fluids. However, the effects of reduced gravity act differently on these two fluids at CHF. The CHF for CFC-113 was decreased in a low gravity level whereas the CHF of water increased. Kato et al.[14] also reported the vanishing of the gravity effect on the heat transfer at high spray volume mass fluxes.

As a follow-up study, Yoshida et al.[15] conducted a more comprehensive study of the effects of gravity on spray cooling. Two different heaters were used in this study. One is similar to the copper block described by Kato et al.[14] except that the surface was plated with chromium and 50 mm in diameter. The other was a glass prism plated with a thin transparent indium tin oxide (ITO) film as the heating element. This transparent glass heater was used in order to visualize the liquid deposition on the heater surface for steady-state spray cooling while a copper block was used for transient spray cooling test. The working fluids used were water and FC-72. A series of ground-based tests and parabolic flight tests were performed by varying the test parameters such as working fluid, heater surface orientation, heat flux at heater surface, the mass flux of the coolant as well as heater types. They reported that gravity level had little effect in the nucleate boiling regime. Moreover, they suggested a coupled effect of gravity and spray volume mass flux on CHF. In the case of a low spray volume flux, neither the magnitude nor the direction of gravity affected CHF. However, the CHF under reduced gravity is higher than that under the hypergravity by 10 percent. They also noted the significant influence of gravity on the film boiling regime when the Webber number was low. And they argued that the deterioration of the heat transfer during the film boiling in the case of low Webber number and reduced gravity condition is due to a lack of secondary impingement on the heater surface.

As indicated by the literature review above, we only found a handful of terrestrial cryogenic spray cooling research papers. However, there has been no attempt on microgravity and reduced gravity cryogenic spray cooling using either room-temperature liquids or cryogens. We believe that the current paper is the first to report the experimental data on cryogenic spray cooling in reduced gravity.

According to transient conduction heat transfer theory[16], if two materials A and B were put together in contact, then the instantaneous heat flux from material A to material B is given by Eq. (1) below,Where are constant temperatures of A and B just before contact, respectively. Also, thermal conductivities of A and B are , respectively. are thermal diffusivities of A and B, respectively. is the interface temperature, while t is the elapsed time after the contact.

We can see that is a function of during the transient[17]. In essence, initially, the heat transfer rate between the two materials is very high, but it also drops off quickly. As a result, for the pulsed flow quenching process, at the moment when the pulsed flow is switched on in a duty cycle, the disk surface gets in contact with a fresh cooling fluid that induces a peak in the heat transfer rate that produces higher cooling rates than the streaming flow case. Based on Eq. (1), the duty cycle (DC) of the pulse flow is the dominant factor on the cooling enhancement solely by the exponential decay of the thermal transport in time, the effect of different periods is only of the second-order effect. Chung et al.[17] found that the pulsed flow would raise the chilldown efficiency up to 67% over the continuous flow case for the convective chilldown of a metal pipe. Chung et al.[17] also reported that the chilldown efficiency increases with decreasing DC, but it is insensitive to the period.

The basis of quenching heat transfer enhancement by the low-thermal conductivity coating is given in Chung et al.[18]. As shown in Chung et al.[19] for convective metal tube chilldown, both thermal diffusivities and thermal conductivities ratios between the stainless steel and the coating material are involved, but clearly, the thermal conductivity ratio dominates the transient process such that the low-thermal conductivity coating can facilitate more than an order of magnitude larger drop of the tube wall surface temperature for the initial period after the quenching is initiated. Chung et al.[18,19] used thin-layers of Teflon for enhancing heat transfer during chilldown of a metal pipe and they found that the coatings could increase the chilldown heat transfer efficiency up to 109 and 176% in terrestrial gravity[19] and microgravity[18], respectively.

The primary objective of the current set of microgravity experiments is to obtain transient quenching heat transfer characteristics of a typical storage tank wall surface simulated by a metal circular disk. The disk transient temperature history or a chilldown curve during spray quenching from room temperature to LN2 saturation temperatures was measured. One of the two disks was coated with a low-thermal conductivity thin Teflon layer to evaluate the heat transfer enhancement. The effectiveness of the coating was evaluated by a comparison of chilldown efficiencies with a coating to those of a bare surface disk. Tests were carried out with a set of pulse flow conditions that includes 40, 50, and 70% duty cycles with 1 s period over a wide range of test section inlet pressure levels and corresponding mass flow rates. The effectiveness of the pulse flow was evaluated by a comparison of chilldown curves with flow pulsations to those with constant and continuous flows.

Methods

Experimental system

Figure 2 shows the photos of the parabolic flight experimental system. The test chamber of this apparatus is designed to comprise two nozzles to spray liquid nitrogen onto two separate stainless-steel circular disks simultaneously. All the system components except the high-pressure gas cylinder fit into a (L × W × H) 1.4 m × 0.8 m × 1 m 8020 aluminum frame. This highly integrated thermal-fluid system was installed on the floor of ZERO-G Corporation’s Boeing 727-200 F aircraft[20] to perform the parabolic flight disk chilldown experiment in a simulated reduced gravity environment. The reduced gravity is achieved through flying the aircraft in a parabolic trajectory and each parabola provides about 17–20 s reduced gravity (10−2g) period.

Fig. 2

Photographic images of the experimental system.

a Front view, b Back view.

The experimental apparatus consists of four essential fluid units together with auxiliary components, fluid piping and instrumentations. The fluid piping schematic and instrumentation diagram is shown in Fig. 3. The 80 L double-wall cryogenic dewar supplies the LN2 to the test section for performing the disk chilldown test as well as provides the LN2 for the prechilling of all the fluid components upstream of the test section prior to the actual chilldown test. Before the experiment, the 80 L dewar is topped off with industrial-grade liquid nitrogen from a standard Airgas 180-Liter dewar through the LN2 fill port. The subcooler is essentially a simple shell-tube heat exchanger and it serves two functions. The first one is to subcool the liquid nitrogen before it enters the test section such that the thermodynamic state of the liquid entering the test section can be determined. During the subcooler operation, the inner finned tube of the subcooler is totally submerged in the liquid nitrogen bath on the shell-side. Since the pressure inside the tube is always higher than that on the shell-side, the liquid nitrogen bath is always colder than that inside the tube. Thus, heat is removed from the liquid in the tube side. The second function is to save the liquid nitrogen during the pre-test chilldown of the upstream components of the test section. The vapor generated on the shell-side is separated by gravity and is vented outside the system.

Fig. 3

The fluid piping schematic and instrumentation diagram of the experimental system.

The valves and important components of the fluid network. Relief valve settings, the burst disk setting, and pressure regulator settings are also included. BD burst disk, BV ball valve, CV check valve, FM flow meter, GN2 gaseous nitrogen, GV globe valve, LN2 liquid nitrogen, PG pressure gauge, PR pressure regulator, PT pressure transducer, RV relief valve, SV solenoid valve, TC thermocouple, Vap vaporizer, 3 V three-way valve.

The test section is basically a vacuum chamber where the cryogenic spray cooling of the disk takes place during the experiment and it is made mostly by off-the-shelf vacuum components. The exploded view of the test section is given in Fig. 4, which shows the configuration of the two test disks and two spray nozzles. Two spray nozzles are placed at the center between two disks inside the 10-inch cubic vacuum chamber. It is noted that the flow direction is perpendicular to the disk and the flow is parallel to the gravity. Two stainless-steel disks, cut from a 16-gauge 304 stainless-steel sheet with a thickness of 0.058 inches, were mounted at opposite sides of the chamber, for example, the front and back or top and bottom. Depending on the purpose of the test, either one set of nozzle and plate or two sets of nozzles and plates are installed. For the ground test, only a single nozzle and one plate were installed inside the test chamber. The orientation of the heat transfer disk surface with respect to the gravity direction is varied by placing the stainless-steel disk at the bottom, side, or top of the cubic test chamber, and they are referred to as upward, vertical, and downward configurations. For the flight tests, two different disk plates and two nozzles were installed as shown in Fig. 4. Each disk is sandwiched by two bored vacuum flanges. Two PTFE gaskets were compressed against the stainless-steel disk such that the cubic test chamber and the back of the disk can be sealed. The outermost flange on each test disk assembly is connected to the vacuum pump such that the back of the stainless-steel disk is insulated from the surroundings by drawing a vacuum to minimize the parasitic heat input from the outside environment. For measuring the disk transient temperature history during the chilldown, 25 thermocouples (TCs) were soldered to the vacuum side (back) of each stainless-steel disc. As shown in Fig. 5, a total of 24 TCs were distributed on the 6 concentric circles (6 rings, R1, R2, …, R6) in addition to one (TC25) placed at the center. Only one TC (TC5) is placed on the outermost point near the outer boundary. TC5 is 3.7″ away from the center of the plate. Eight MINCO film heaters (Hap 6945 and 6946) are attached to the back of the disk to reheat the stainless-steel plate back to the initial temperature for the next test after each chilldown test.

Fig. 4

A CAD drawing of the test section.

The test chamber is 10-inch cubic and it houses two spray nozzles.

Fig. 5

Schematic of thermal couple placement.

Locations of 25 thermal couples are shown on six concentric circles.

The flow coming out of the test section goes into two separate vaporizers in parallel (labeled as Vap2 and Vap3 in Fig. 3). The vaporizers are basically heat exchangers made from tube bundles, which evaporate any remaining liquid nitrogen coming out of the test section and also heat up the cold nitrogen vapor to above 0 °C before venting pure vapor out of the system. Each vaporizer is made by packing eight grooved copper tubes that have star-shaped inner insertions into a 2.5″ schedule 40 stainless-steel pipe. The tube bundles are heated up to 200 °C before each test by a high-temperature heating tap that is wrapped around the outer surface of the stainless-steel pipe. The Labview program monitors and controls the on and off of the heating tap by the combination of a K-type thermocouple, NI 9211 thermocouple input module, NI 9472 digital output module, and a mechanical relay. If the experiment is performed onboard the aircraft, the gas coming out of the vaporizers is vented outside the aircraft cabin through rubber hoses that connect to the vent ports on the cabin wall. Similarly, another vaporizer, Vap1 ensures the proper venting of gaseous nitrogen coming out of the subcooler.

The data acquisition system including the Labview program and National Instrument Compact DAQ hardware collects all sensor data and displays the real time on a laptop at a sampling rate of 16 Hz. NI 9214 TC modules read all the T-type TCs (Omega). NI 9205, an analog input module, reads all the voltage signals from pressure transducers (Omega PX 409V5A) and the 4–20 mA current signals (through a 249-ohm resistor) from the Coriolis liquid flow meter (Micro motion CMF025). The Labview program controls the opening and closing of the solenoidal valve, SV1, through a combination of NI USB 6009 and a Solid-State relay. In the case of a continuous spray, the relay energizes the solenoid valve after receiving a continuous voltage signal. However, in the case of an intermittent spray, the relay energizes and de-energizes SV1 according to a rectangular waveform voltage signal from the VI.

In the current experiment, we used two types of disks. In addition to the bare surface stainless-steel disk, we also added a coated disk that is a stainless-steel disk coated with a low-thermal conductivity thin-film Teflon layer on one side of the disk surface. Specifically, the coating material was made of Fluorinated Ethylene Propylene (FEP) by DuPont and classified by DuPont as Teflon 959G-203 that is a black color paint and has a thermal conductivity of 0.195 W/mK (DuPont publication[21]). The coating was put on using the dip and drain process. The thickness of the coating is estimated at around 20–30 microns. The thickness of the Teflon coating was estimated by previous experiences obtained from an identical coating method. In our previous pipe chilldown experiment[19], the thickness of the coating on the tube inner surface was measured by the high-resolution X-ray computer tomography (CT) scan. Since we used the same method to coat the disk plate as that used in the tube and expected the thickness of the coating is similar to that of the tube coating.

Experimental procedure

To perform the chilldown test, mainly four steps are followed, and these are initial starting, precooling, testing, and reheating. The initial starting is the step where all the electrical devices are turned on. This includes running the preprogrammed Labview script and turning on the vacuum pump (Turbo Lab 80). Once the Labview script is running, it will automatically set the output voltage of the DC power supply for the pressure transducers and turn on the heating cables of the vaporizers. This step takes about 30 min mainly due to the time required for heating up the vaporizers to 200 °C. Meanwhile, the globe valves, GV1, GV2, GV3 are open for the next step. The second step involves the precooling and prechilling of all the piping and fluid components upstream of the test section to make sure that only the liquid phase of working fluid enters the test section and it proceeds first by rotating the three-way valve, 3V1, from GV1 to GV2, and setting the desired testing pressure on the pressure regulator, PR. Once the solenoidal valve SV2 is open by clicking the virtual button on the Laptop screen, then the liquid nitrogen starts to flow from the 80 L dewar into the shell-side of the subcooler. Before the liquid nitrogen can fill up the shell-side of the subcooler, the flow path upstream of the test section has to be chilled down. This step prevents the boil-off of liquid nitrogen before it flows into the test section. Once the inner tube of the subcooler is full (can be determined by the profile reading of the TC located inside the subcooler), the system is ready for the experiment. The chilldown test is started simply by clicking the virtual start bottom on the laptop screen, then the solenoid valve, SV1, will be opened according to the preset waveform signals either to continuously or intermittently flow nitrogen into the test section for spraying on the target disks. Once all the temperature readings from the TCs drop to the liquid nitrogen temperature and maintain at a steady-state, the disk chilldown experiment is complete. The solenoid valve, SV1, is then closed by clicking the virtual stop bottom. Next, the reheating step starts to prepare the stainless-steel disc plates for the next test. To heat up the plates after chilldown, the film heaters are turned on by clicking the heating virtual bottom on the screen. Once the plates are heated back to room temperature, the film heaters are turned off. This marks the beginning of the next cycle of testing which starts with the precooling step.

In addition to the experimental procedure discussed above, next, we need to mention the simulated microgravity environment on the parabolic flight. The variable gravity condition inside the airplane was created when the airplane is flying a parabolic trajectory[20]. The microgravity period is always sandwiched between two 1.8-g periods where g is the earth’s gravity of 9.81 m/s2. The microgravity period nominally lasts between 18-25 s. For our research flights, in order to maintain the acceleration levels within ±0.01 g, the microgravity period is around 18–20 s.

Experimental uncertainty

Table 1 lists all the uncertainties or the independently measured quantities. The uncertainty for the chilldown thermal efficiency is discussed above in the Results section.

Table 1

Measured quantities and their uncertainties.

Symbol	Quantity	Measurement method	Uncertainty
T	Plate temperature	T-type thermocouple	1 K or 1.5% below 273 K
Pc	Test section pressure	Kulite CTL-190M140BARA	7 kPa
δ	SS plate thickness	Calipers	0.03 mm
Rx	Radial position of TCs on Ring x	Ruler	1.6 mm (1/16″)
ṁ	LN2 mass flow rate	Coriolis flow meter	0.3%
M	Mass of the fluid components (tube, tee, nozzle)	scale	0.1 g

Table 2

Experimental conditions for the six flight cases.

Case	Pin (psig)	Duty cycle (%)	Period (second)	g-level
1	80	100	NA	microgravity
2	80	40	1	microgravity
3	80	70	1	microgravity
4	60	100	NA	microgravity
5	90	100	NA	microgravity
6	90	50	1	microgravity

Table 3

Individual uncertainty of the independently measured quantities.

Symbol	Quantity	Measurement method	Uncertainty
D	Diameter of the target area	Rulers	1.6 × 10−3 m
ρ	Density of test section	NIST website	1%
cp	Specific heat capacity of the test section	NIST website	5%
ΔT	Temperature difference between the initial and the end of the test	T-type thermocouple	1.5%
τ	Thickness of test section	Calipers	3 × 105 m
hfg	Latent heat of nitrogen	NIST website	5%
m˙_l	Mass flow rate of liquid nitrogen	Coriolis flow meter	0.3%

Table 4

Chilldown efficiencies and mass flow rates for the six flight cases.

	Reduced-G			Ground-G
Case	Teflon coated	Bare SS304	Mass flow rate, kg/s	Teflon coated	SS304
1 (micro-g) 80 psig, continuous flow	18.54%	13.38%	0.0253	23.19%	19.54%
2 (micro-g) 80 psig 40% DC, 1 s Period	28.50%	21.26%	0.0350	28.49%	21.74%
3 (micro-g) 80 psig 70% DC, 1 s Period	23.48%	13.65%	0.0239	21.17%	17.78%
4 (micro-g) 60 psig, continuous flow	28.40%	17.83%	0.0100	24.50%	18.58%
5 (micro-g) 90 psig, continuous flow	16.87%*	12.96%*	0.0264*	17.26%	14.32%
6 (micro-g) 90 psig 50% DC, 1 s Period	23.88%*	18.63%*	0.0255*	20.52%	16.76%

Table 5

Improvement of chilldown efficiency by coating.

	Reduced-G	Ground-G
Case	Percent improvement	Percent improvement
1 (micro-g) 80 psig, continuous flow	38.6%	18.7%
2 (micro-g) 80 psig 40% DC, 1 s Period	34.1%	31.1%
3 (micro-g) 80 psig 70% DC, 1 s Period	72.0%	19.1%
4 (micro-g) 60 psig, continuous flow	59.3%	31.9%
5 (micro-g) 90 psig, continuous flow	30.2%	20.5%
6 (micro-g) 90 psig 50% DC, 1 s Period	28.2%	22.4%