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Nitrogen flow boiling and chilldown experiments in microgravity using pulse flow and low-thermally conductive coatings

Nitrogen flow boiling and chilldown experiments in microgravity using pulse flow and... www.nature.com/npjmgrav ARTICLE OPEN Nitrogen flow boiling and chilldown experiments in microgravity using pulse flow and low-thermally conductive coatings 1 2 2 2 2 3 3 4 Jason Hartwig , J. N. Chung , Jun Dong , Bo Han , Hao Wang , Samuel Darr , Matthew Taliaferro , Shreykumar Jain and Michael Doherty The enabling of in-space cryogenic engines and cryogenic fuel depots for future manned and robotic space exploration missions begins with technology development of advanced cryogenic fluid management systems upstream in the propellant feed system. Before single-phase liquid can flow to the engine or customer spacecraft receiver tank, the connecting transfer line must first be chilled down to cryogenic temperatures. The most direct and simplest method to quench the line is to use the cold propellant itself. When a cryogenic fluid is introduced into a warm transfer system, two-phase flow quenching ensues. While boiling is well known to be a highly efficient mode of heat transfer, previous work has shown this efficiency is lowered in reduced gravity. Due to the projected cost of launching and storing cryogens in space, it is desired to perform this chilldown process using the least amount of propellant possible, especially given the desire for reusable systems and thus multiple transfers. This paper presents an assessment of two revolutionary new performance enhancements that reduce the amount of propellant consumed during chilldown while in a microgravity environment. Twenty-eight cryogenic transfer line chilldown experiments were performed onboard four parabolic flights to examine the independent as well as combined effect of using low thermally conductive coatings and pulse flow on the chilldown process. Across a range of Reynolds numbers, results show the combination significantly enhances performance in microgravity, with a reduction in consumed mass up to 75% relative to continuous flow for a bare transfer line. npj Microgravity (2022) 8:33 ; https://doi.org/10.1038/s41526-022-00220-9 INTRODUCTION volume fill fractions in the customer receiver tank, and most in- space engines require single-phase liquid up to the injectors. The enabling of in-space cryogenic engines and cryogenic fuel Therefore, without advanced CFM technologies upstream in the depots for future manned and robotic space exploration missions feed system and storage tank, vapor ingestion is inevitable, which begins with technology development of advanced cryogenic fluid can lead to combustion instabilities within the engine. Further management (CFM) systems upstream in the propellant feed exacerbating the transfer process in microgravity is the unknown system. Cryogenic propellants offer significantly higher perfor- location of the liquid and vapor phases in the tank as well as mance relative to storable counterparts, such as hydrazine, owing reduced heat transfer. to a higher specific impulse and higher energy density. Further, Before single-phase liquid can flow to the engine or customer safety and environmental concerns over the use of toxic storable spacecraft receiver tank, the connecting transfer line must first be propellants have led to the ongoing examination of more “green” chilled down to cryogenic temperatures. Chilldown, or quenching, propellants such as liquid methane as alternate fuel sources. Aside is defined as the transient process of cooling hardware down to from nuclear thermal propulsion systems , no other known pure cryogenic temperatures so that vapor-free liquid can eventually chemical propulsion system propellant combination can deliver a flow between two points of interest. The most direct and simplest higher ISP than liquid hydrogen/liquid oxygen. However, there are method to quench the line is to use the cold propellant itself. challenging aspects when working with cryogens due to inherent When a cryogenic fluid is initially transferred through a system, thermo-physical properties. Particularly for the current work, the the tube walls and hardware (e.g. valves) undergo a transient low normal boiling point (NBP), low surface tension, and high chilldown prior to reaching a steady state of operation. Chilldown susceptibility to parasitic heat leak leads to unwanted boiling and thus involves unsteady two-phase heat and mass transfer and flow two-phase flow during propellant transfer. boiling. While boiling is well known to be a highly efficient mode 2,3 Cryogenic fuel depots ,defined as an Earth-orbiting propellant of heat transfer, previous work has shown this efficiency is storage vessel that would house cryogenic propellant to allow significantly lowered in reduced gravity, both for room tempera- 4–8 spacecraft to refuel, have four stages: (1) acquisition of the storage ture fluids as well as cryogens . Due to the projected cost of tank liquid, (2) chilldown of the connecting transfer line hardware, launching and storing cryogens in space, it is desired to perform (3) chilldown of the receiver tank, and (4) fill of the receiver tank, this chilldown process using the least amount of propellant as all in the microgravity of space; this paper focuses on the second possible, especially given the drive towards reusable systems and stage, chilldown of the transfer line. Meanwhile, cryogenic engines thus multiple transfers. also require acquisition of the storage tank liquid and chilldown of Numerous cryogenic flow boiling quenching experiments have the transfer line. Cryogenic fuel depots will require very high liquid previously been conducted on bare tubes, the results up to 2018 1 2 3 4 NASA Glenn Research Center, Cleveland, OH 44135, USA. University of Florida, Gainesville, FL 32611, USA. The Aerospace Corporation, El Segundo, CA 90245, USA. Georgia Tech University, Atlanta, GA 30332, USA. email: Jason.W.Hartwig@nasa.gov Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA 1234567890():,; J. Hartwig et al. of which are summarized in ref. . Key contributions in 1-g were conductive coatings and pulse flow on the chilldown process. 4,5,10–23 provided by , which investigated the effect of mass flux, While previous experiments have reported the effects of pulse inlet state, pressure, and flow direction on cryogenic tube flow and coatings on transfer line chilldown in Earth-gravity, this is chilldown, predominately using liquid nitrogen (LN ) and liquid the first report of pulse flow and the combined effect of pulse flow hydrogen (LH ). Since 2018, five more cryogenic quenching with a coated tube in a microgravity environment. studies on bare tubes that passed the data filtering criteria from ref. by Jin et al. have been added to the cryogenic database to 24 25 RESULTS AND DISCUSSION cover low Reynolds (Re) number LN , liquid argon , and liquid 26 27,28 oxygen chilldown experiments while added high Re number Test matrix chilldown tests with LH . Hartwig et al. recently summarized Table 1 lists the complete flight test matrix. Ground tests were cryogenic quenching flow boiling trends over the consolidated performed at the University of Florida, while flight tests were literature, across multiple flow regimes, mass fluxes, inlet states, conducted during the low-gravity portion of the classic parabolic and gravity levels. For all cryogens, the chilldown process is highly trajectory followed by the flight provider ZeroG. Pressure is the dominated by the film boiling regime for bare tubes (for measured pressure at the inlet to the test section, time-averaged quantum fluids such as hydrogen and helium, there are additional over the test duration. Period is the sum of valve “on” and “off” time factors at play). When a cryogen is introduced into a warm tube, for a pulse flow test cycle. The duty cycle is the ratio of the valve “on” especially at high mass flux and low inlet equilibrium quality, a time to the period. For example, with a period of 3 s and a duty cycle vapor film blanket surrounds the liquid core which acts as an of 10%, the valve is on 0.3 s and off for 2.7 s. G level is the gravity insulator that inhibits heat transfer between cold liquid and warm level as read by accelerometers attached to the experimental rig tube. At lower mass flux and saturated inlet states, dryout occurs while on the flight. Note that a few of the flight tests were conducted over a longer distance along the tube as in the case of traditional at a g-level higher than nominal; these were deemed “Martian gravity fluids . Film boiling heat transfer is a highly inefficient process tests”. Coating thickness in number of layers, “L”, is described in the relative to transition and nucleate boiling. In most instances, film “Methods” section. boiling can persist for >85% of the total time needed to chill the Chilldown time was determined as follows: In practice, the most tube down to the saturation temperature of the cryogen. Once the stringent chilldown criteria would be determined from a Leidenfrost point is reached, chilldown proceeds into transition measured stream temperature downstream of the test section boiling, nucleate boiling, and then single-phase liquid convective reading lower than the saturation temperature based on the flow. In microgravity, this poor heat transfer is exacerbated by the downstream pressure; however, this measurement was not lack of buoyancy force; cryogenic film boiling heat transfer was available for the current tests. Based on boiling heat transfer shown to be 25% lower at low to modest Re flows relative to 1-g . theory, nucleate boiling would end when the inner surface At very high Re, inertial forces can overcome gravitational forces temperature drops below that of the onset-of-nucleate boiling such that gravity no longer affects flow boiling (although this (ONB). As a result, the wall heat flux would switch from higher has not been demonstrated yet for cryogens). boiling heat flux to much lower single-phase convective heat flux To overcome this hurdle in poor performance, researchers have that would reflect a change on the outer wall surface temperature recently investigated low thermally conductive materials applied gradient with time. A computed inner wall temperature could also to the inner tube walls and the effect of such coatings on the not be used to determine end of chilldown; while the inverse chilldown process. The coating acts as an insulator between the conduction method of Burggraf can be used to determine inner cold propellant and warm wall, resulting in an inner wall surface wall temperature for bare tubes, due to the unknown thermal temperature that reaches the Leidenfrost point without cooling contact resistances between the coated layer and tube inner wall the entire tube mass. Recent 1-g experiments conducted in the 33,34 35 as well as among adjacent coated layers, inner wall temperature United States and China independently confirmed that a could not be determined for coated tubes. Therefore, outer wall Teflon coated tube could reduce chilldown times up to 75% over temperature data had to be used to determine end of chilldown. an uncoated stainless steel (SS) tube using LN . Both researchers Three chilldown criteria were explored: (1) the averaged exit outer also investigated the effect of the Teflon coating thickness on wall temperature was compared to the liquid saturation chilldown performance, and both showed that thicker coatings temperature (based on local downstream pressure), (2) the first led to faster chilldown times. However, as the coating thickness derivative of outer wall temperature (with respect to time) increased further there was an apparent point of diminishing reaching and remaining near 0 K/s (due to minimal convective returns because the chilldown curves (wall temperature versus heat transfer between single-phase liquid and tube), and (3) a time) converged at the highest tested thicknesses. Coated tubes peak value in the second derivative of outer wall temperature offer hope to combat the intrinsically poor film boiling heat (with respect to time) which would indicate the slope change in transfer in microgravity . the chilldown curve occurring at onset of nucleate boiling (ONB). A second way to enhance poor chilldown performance is to use 37 38 The first method was found to be unreliable due to inaccurate pulse flow. Demonstrated using both LH and LN , in pulse 2 2 chilldown time estimations attributed to the significant difference flow, the inlet valve is cyclically opened and closed with a between inner and outer wall temperature at higher layers of specified duty cycle (DC) and pulse width until the desired degree coating. The third method also yielded inaccurate chilldown time of chilldown is reached. The advantage of pulse flow is lower mass estimations attributed to the absence of a true global maxima in consumption over traditional continuous flow due to more the second derivative at higher layers of coating. Therefore, the efficient usage of latent and sensible energy of the fluid, with second method using the first derivative (typically using test the disadvantage being potential valve fatigue and/or failure and section averaged temperature), slightly conservative but consis- added complexity in operation. tent across all scenarios, was used to determine the end of The purpose of this paper is to present an assessment of two chilldown in all test cases. new performance enhancements that reduce the amount of Chilldown mass was the total consumed LN mass at the end of propellant consumed during chilldown while in a microgravity 2 chilldown as read by the flow meter downstream of the test environment, and to investigate if the mass savings holds in section: microgravity. Twenty-eight LN transfer line chilldown experi- ments were performed onboard a parabolic flight that simulated Z end space microgravity conditions to examine the independent as well m ¼ mdt _ (1) LN2 as combined performance gains of using low thermally 0 npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA 1234567890():,; J. Hartwig et al. Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA npj Microgravity (2022) 33 Table 1. Flight test matrix. # Test name Inlet pressure Period [s] Duty G Level [g/g ] Coating Chilldown Total chilldown Steady-state Re Steady-state mass cycle [%] thickness time [s] mass [kg] flux [kg/m -s] F1 Flight_day1_1 550 kPa [80 psia] – 100 0.38 Bare 15.1 1.08 71,976 548.87 F2 Flight_day1_2 550 kPa [80 psia] – 100 0.05 Bare 14.6 0.872 67,324 533.22 F3 Flight_day1_3 550 kPa [80 psia] 2 10 0.05 Bare 16 0.597 53,859 426.58 F4 Flight_day1_4 550 kPa [80 psia] 3 10 0.05 Bare 16.7 0.614 64,092 507.63 F5 Flight_day1_5 340 kPa [50 psia] – 100 0.05 Bare 70.6 0.264 1761 22.97 F6 Flight_day1_6 340 kPa [50 psia] 4 10 0.05 Bare 78.1 0.279 1383 18.04 F7 Flight_day2_1 550 kPa [80 psia] 3 7 0.38 Bare 21.3 0.637 32,621 347.47 F8 Flight_day2_2 340 kPa [50 psia] – 100 0.05 Bare 80.3 0.279 1566 20.42 F9 Flight_day2_3 340 kPa [50 psia] 5 4 0.05 Bare 162 0.304 881 11.49 F10 Flight_day2_4 340 kPa [50 psia] – 100 0.05 Bare 172 0.16 992 15.94 F11 Flight_day2_5 440 kPa [65 psia] 3 7 0.05 Bare 21.3 0.645 29,761 306.99 F12 Flight_day2_6 440 kPa [65 psia] – 100 0.05 Bare 21.1 1.174 53,021 459.65 F13 Flight_day3_1 550 kPa [80 psia] – 100 0.05 7 L 8.19 0.469 63,912 506.19 F14 Flight_day3_2 550 kPa [80 psia] 3 7 0.05 7 L 9.94 0.339 38,308 351.49 F15 Flight_day3_3 550 kPa [80 psia] – 100 0.05 7 L 6.69 0.309 45,760 340.84 F16 Flight_day3_4 550 kPa [80 psia] 3 7 0.05 7 L 10 0.248 43,178 329.21 F17 Flight_day3_5 340 kPa [50 psia] – 100 0.05 7 L 12.8 0.0881 8945 73.36 F18 Flight_day3_6 440 kPa [65 psia] 3 7 0.05 7 L 9.31 0.283 38,265 364.8 F19 Flight_day3_7 440 kPa [65 psia] – 100 0.05 7 L 7.06 0.514 83,353 649.18 F20 Flight_day3_8 440 kPa [65 psia] – 100 0.05 7 L 7.88 0.585 88,127 682.68 F21 Flight_day3_9 340 kPa [50 psia] – 100 0.05 7 L 12.8 0.0855 10,949 90.82 F22 Flight_day4_1 550 kPa [80 psia] – 100 0.05 4 L 4.69 0.281 74,196 558.16 F23 Flight_day4_2 550 kPa [80 psia] 3 7 0.05 4 L 9.06 0.269 46,373 348.85 F24 Flight_day4_3 550 kPa [80 psia] 2 10 0.05 4 L 6.19 0.237 63,067 474.44 F25 Flight_day4_4 440 kPa [65 psia] – 100 0.05 4 L 4.5 0.23 59,931 476.14 F26 Flight_day4_5 440 kPa [65 psia] 3 7 0.05 4 L 10 0.288 50,941 404.72 F27 Flight_day4_6 340 kPa [50 psia] – 100 0.05 4 L 7.81 0.0555 8061 66.22 F28 Flight_day4_7 550 kPa [80 psia] 3 10 0.05 4 L 7.25 0.274 64,048 500.43 These cases were conducted over multi-parabolas; the calculations also include the high G part. Airplane accelerated to 1 g before end of chilldown. The initial temperature is lower than other cases because there was not enough time to reheat. Missing time-series mass flow rate data. J. Hartwig et al. Fig. 1 Test F2 (Flight_Day1_2, continuous flow, 550 kPa source pressure, 67324 Re, bare tube surface, 0.05 G level (g/g0)). a Chilldown curve based on average exit wall temperature, b Boiling curve based on average exit wall temperature, c Chilldown curve of all TCs. where t is the end of chilldown time and mðtÞ is the time- Averaging was done by adding the temperatures and dividing by end dependent LN mass flow rate measured by the gas flow meter. the number of sensors. Figure 1c illustrates the chilldown curve of Steady state Reynolds (Re) number (defined at the end of all thermocouples (TCs) placed on the tube outer wall according to chilldown when single phase liquid flow is established) and mass Fig. 5d, e in the “Methods” section. Errors bars are plotted but flux were evaluated using inner diameter and saturation condi- barely discernable. Three boiling regimes, film boiling (FB), tions based on the measured test section pressure: transition boiling (TB), and nucleate boiling (NB), and single- phase convection are separated by three critical points, the 4m _ Re ¼ (2) Leidenfrost Point (LFP), Critical Heat Flux (CHF), and the onset of πDμ nucleate boiling (ONB). The chilldown curve begins in the film 39 boiling regime where the cold liquid entering the warm tube For uncoated tubes, the method of Burgraff was used to experiences violent boiling. Depending on the local conditions, determine inner wall temperature and transient radial heat the flow will proceed into dispersed flow FB (high quality, low conduction through the tube as follows: subcooling, low mass flux) or inverted annular FB (low quality, 2 2 2 3 4 2 r r ðÞ ρc r r r r 00 o dT P i r d T 40,41 i o i o o i o high subcooling, high mass flux) . The high wall surface q ¼ ρc þ   ln P 2 2r dt k 16 16r 4 r dt i i o temperature causes the liquid to completely vaporize before 3 5 4 2 3 6 2 3 4 r 3r r r r ðÞ ρc 3r r r r r d T P i o i o i o o i ri o i ri o þ  þ   ln  ln reaching the surface resulting in an inner liquid core and outer 2 3 k 384 128 128 384r 128 r 32 r dt i o o annular vapor core. This vapor blanket along the wall insulates the (3) warm pipe from the cold liquid, causing the temperature of the where q is the radial heat flux through the tube, ρ, C , and k are pipe to decrease, albeit slowly. FB is the least efficient quenching the tube density, specific heat, and thermal conductivity, mechanism. As the transfer line chills down, the system respectively, r and r are the inner and outer radii, and T is the approaches the LFP, or rewet temperature, where heat flux is at i o o outer wall temperature. Heat transfer coefficient was then a minimum (during boiling). Heat transfer here is a minimum due computed as follows: to the inefficiency of heat transfer between cold vapor and wall. LFP is also characterized by the onset of a rapid drop in wall 00 00 o 00 00 00 q þ q þ q þ q þ q axial r rad solidcond gascond temperature. As shown in Fig. 1c, the LFP occurs at later times for (4) h ¼ quench TCs located farther downstream. This trend demonstrates the T  T i sat location of the quenching front as it propagates downstream as where q is the axial conduction along the tube; the terms in axial chilldown evolves. The flow then proceeds and passes quickly parenthesis are the radiation, solid conduction, and gaseous through TB, characterized by intermittent liquid contact along the conduction parasitic heat leak terms, respectively, T is the inner walls. TB ends when liquid is in full contact with the walls at the wall temperature which comes from Burgraff’s method, and T is sat point of CHF. Heat transfer is a maximum at CHF due to the highly the saturation temperature based on the measured pressure. The efficient cooling process of boiling. Nucleate boiling follows, method to calculate the different heat fluxes in Eq. 4 has been where heat is transferred by vapor bubbles formed in surface 22 27 shown in many other papers, see for example ref. and ref. . cavities that are swept away from the tube surface. Depending on the inlet conditions, NB can be liquid-convection dominate or Governing physics of chilldown nucleation-dominate . As the wall cools further, the tube inner Figure 1a shows the chilldown curve of averaged exit wall surface approaches the ONB, characterized as the point at which temperature (TC5, TC10, TC15) and Fig. 1b shows the boiling curve the system evolves from nucleate two-phase cooling to single- based on the averaged exit wall temperature in microgravity. phase liquid convection and an obvious slope change in the npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA J. Hartwig et al. Fig. 2 Effect of teflon coating thickness in microgravity: 550 kPa source pressure, continuous flow, 0.05 G level (g/g0). F2 (Flight_Day1_2, 67324 Re, bare surface) versus F22 (Flight_Day4_1, 74196 Re, 4 L coating) versus F13 (Flight_Day3_1, 63912 Re, 7 L coating) a Chilldown curve, b exit pressure, c mass flux, and d total propellant mass consumed based on averaged exit wall temperature. chilldown curve. Vapor-free liquid marks the end of the chilldown wall temperature for the coated tubes will be significantly lower test. The single-phase cooling causes the wall temperature to drop since the coating restricts the heat transfer between inner and slowly to the liquid saturation temperature and then remain outer walls. Second, less mass is consumed for coated over bare steady as heat transfer reduces to near zero. In microgravity, tubes as substantiated in Fig. 2d; the 4 L and 7 L coated cases have circumferential TCs at each station have almost identical chilldown 68 and 46% propellant mass savings over the bare tube, behavior at any axial distance from the inlet; stratification effects respectively. normally seen for horizontal tubes in 1-g disappear, leading to Third, however, there is an apparent point of diminishing axisymmetric flow patterns through the tube, and thus uniform returns; this trend of improved chilldown performance upon chilldown circumferentially. addition of coating is reversed when the number of coating layers is increased from 4 to 7 because the chilldown time is faster for 4 L (4.7 s) compared to 7 L (8.2 s) case. Similarly, from 4 L to 7 L, the Bare vs. coated tube in microgravity, continuous flow propellant mass savings and chilldown efficiency are reduced. This Figure 2a–d plot chilldown curves, exit pressure, mass flux, and crossover in performance and possible existence of an optimal total consumed liquid mass for bare, 4 L, and 7 L coated tubes for coating layer is explained by counteracting heat transfer higher steady state Re (63,912–74,196). The initial fluctuations in mechanisms: (1) the low thermal conductivity of the coating layer pressure measurements in Fig. 2b are due to the transient nature facilitates the faster temperature drop of tube inner surface by of the flow at start of the test. Shortly after the transient start, restricting heat transfer between inner surface and bulk of the downstream pressure measurements reach their steady-state metal tube and (2) the low thermal conductivity coating also value and remain there until at least the end of chilldown in all creates a thermal resistance that restricts the heat conduction three cases. The mass flux of the 4 L coating case in Fig. 2cisa between bulk of the tube and cooling fluid. With these contrasting straight horizontal line because of missing timed mass flow rate mechanisms at play, the thickness of the coating must be such data for that run; a linear correlation was developed between that it is thick enough to quickly lower the tube inner surface averaged inlet pressure and averaged mass flow rate for cases temperature while being thin enough to facilitate fast wall with available mass flow rate data that were run at 0.05 g level and chilldown. However, the presence of the coating accelerates were completed under one parabola. This linear correlation was chilldown as evident in any comparison between bare and coated then used to calculate an average mass flow rate for cases with tube at similar thermodynamic conditions. missing mass flow rate data (but available inlet pressure data). Trends are as follows: First, coating the inner wall of the tube Continuous versus pulse flow in microgravity, bare tube drastically affects the chilldown behavior and leads to faster chilldown times. The low thermally conductive Teflon layer acts as Figure 3a–e plot chilldown curves, heat transfer coefficient, an insulator between cold fluid and warm wall; the inner surface pressure, mass flux, and total consumed liquid mass for temperature chills down quickly without cooling the entire tube continuous flow and pulsed flow at a period of 2 s and duty mass. The lower inner wall surface temperature earlier on means cycle 10% (valve on 0.2 s, valve off 1.8 s) and for period 3 s and that the Leidenfrost point is reached faster such that the liquid can duty cycle 10% (valve on 0.3 s, valve off 2.7 s) at higher Re stay in contact with the tube for the heat transfer to be in TB and (53859–64092). Trends are as follows: First, both continuous and NB that reduces the poor heat transfer film boiling time; this is pulse flow exhibit the same chilldown curve and proceed through substantiated by the drastic slope change for 4 L and 7 L tube the same transition points. For pulse flow, the longer the valve-off indicating the LFP is reached earlier on relative to the bare tube. time, the more the tube temperature stabilizes as residual cooling Note that Fig. 2a plots outer wall temperature; the actual inner due to blowdown diminishes. Second, from Fig. 3a, e, it is clear Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA npj Microgravity (2022) 33 J. Hartwig et al. Fig. 3 Effect of pulse flow on a bare tube in microgravity: 550 kPa source pressure, bare tube surface, 0.05 G level (g/g0)—F2 (Flight_Day1_2, 67324 Re, continuous) versus F3 (Flight_Day1_3, 53859 Re, pulse 2 s 10%) versus F4 (Flight_Day1_4, 64092 Re, pulse 3 s 10%). a Chilldown curve, b Heat transfer coefficient versus wall superheat, c Exit pressure, d mass flux, and e total propellant mass consumed based on averaged exit wall temperature. Fig. 4 Combined effect of coatings and pulse flow in microgravity: 0.05 G level (g/g0)—F2 (Flight_Day1_2, 550 kPa source pressure, 67,324 Re, bare surface, continuous) versus F28 (Flight_Day4_7, 550 kPa source pressure, 64,048 Re, 4 L Coating, 3 s 10%). a Chilldown curve and b total propellant mass consumed based on averaged exit wall temperature. that pulse flow achieves chilldown using less propellant but at the Table 1, there is 29–32% mass savings with pulse flow in cost of longer chilldown time due to better use of sensible and comparison to continuous flow at these flight conditions. Third, latent energy of the fluid. Figure 3c, d shows fluctuations in for a fixed duty cycle, reducing the valve-open time leads to pressure and mass flux that are due to valve cycling, that these slightly shorter chilldown times (although not shown directly in fluctuations continue until end of chilldown, and that the Fig. 3) and, slightly less propellant consumption as shown in fluctuation amplitudes are higher for longer periods. From Fig. 3a, e; this trend compares well with previous pulse flow tests npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA J. Hartwig et al. Fig. 5 Experimental Design. a Flight system piping and instrumentation diagram, b actual flight line chilldown rig, c pre-cooler, d test section, vacuum chamber, and thermocouple locations, A: Inlet, B: 3.81 cm (1.5 in) long tube section, C: ultra-torr fitting, D: left flange, E: right flange and e thermocouple locations at each station. 33 37 for both LN and LH . Fourth, for this particular comparison, METHODS 2 2 Fig. 3b shows that continuous flow exhibited a higher CHF over Experimental description pulse flow, and that reducing the valve open time reduced the The authors have completed four successful cryogenic line CHF. Because of the temperature stabilization when the valve was chilldown parabolic flight campaigns including the current cycled off, the temperature does not drop as rapidly in pulse campaign between 2015 and 2020 and are familiar with system compared to continuous flow which caused the wall temperature designs, troubleshooting, issues, and failures that arise with first derivative term to be lower at CHF in pulse flow. However, if microgravity flight testing. The fourth-generation system was 8 36 the CHF was traversed when the valve was on, it is expected that modified based on flights from the first- and second-generation the pulse flow heat transfer coefficient (HTC) would be nearly systems. As before, the system is intended for both ground and equivalent to that of continuous flow. For bare tubes in flight experiments. Figure 5a shows a system flow network and piping and instrumentation diagram while Fig. 5b shows a picture microgravity, while higher frequency, shorter pulse widths are of the actual flight rig. favorable from a chilldown efficiency standpoint, more valve LN was supplied to the system from an 80-liter vacuum- cycles implies higher risk of valve degradation and potential jacketed dewar, with a relief valve set at 861 kPa. A gaseous failure. Therefore, there is an inherent trade-off in which the nitrogen (GN ) cylinder initially pressurized at 15 MPa was used to optimal valve duty cycle could be determined. pressurize the dewar to a set value for each test, which ranged between 90 and 830 kPa absolute pressure. Dewar pressure was managed by a pressure regulator that controlled the dewar Performance gain of combined pulsed flow and coated tubes pressure to within 35 kPa of the set value during each test. in microgravity Depressurization was carried out by opening the globe valve 2 Figure 4a, b plot chilldown curves and total consumed liquid mass (GV2) and the three-way ball valve 1 (3V1) to allow ullage gas to for bare tube with continuous flow and 4 L coated tube with vent to the atmosphere. pulsed flow characterized by a period of 3 sec and duty cycle of The dewar was used to supply the LN both for prechilling the 10% (valve on 0.3 s, valve off 2.7 s) at higher Re (64,048–67,324). plumbing upstream of the test section and for conducting the actual chilldown experiment. LN was delivered through valve Trends are as follows: First, the effect of coating on reducing GV3 that was connected through a 1.2 m long, 1.27 cm outer chilldown time seems to outweigh the effect of pulse flow on diameter (OD), 1.18 cm inner diameter (ID) 304 (stainless steel) SS increasing chilldown time as evidenced by the sharp drop in braided hose to the inner tube of the precooler (or subcooler) temperature at ~4 s in Fig. 4a for the coated tube. Second, the shell-tube heat exchanger shown in Fig. 5c. The subcooler served individual benefits of propellant mass savings with coating and three purposes: (1) to preserve subcooling of the liquid from the pulse flow are nearly perfectly superimposed, leading to a 76% storage tank flowing through the transfer line by eliminating reduction in propellant consumption. Results thus show that high parasitic heat leak, (2) to slightly subcool the LN in the transfer performance is still achieved in microgravity for pulse flow with a line, since the saturation temperature of the shell side was always low thermal conductivity coating which leads to a reduction in lower than the tube side, and most importantly (3) to ensure chilldown time and mass and increase in chilldown efficiency over single-phase liquid at the inlet of the test section. The liquid level continuous flow with a bare tube. of the nitrogen pool was monitored by three thermocouples (TC) Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA npj Microgravity (2022) 33 J. Hartwig et al. inside the subcooler, two at the shell-side and one at the outlet to x-ray scans of the tube cross sections were obtained using a Vap1. The temperature readings of these TCs were displayed on a Phoenix v|tome|x M system in the Nano Research Facility at the laptop in real-time. The level of the LN pool was inferred from the University of Florida. Scanning was carried out using a 240 kV TC insertion depth. A 2.5 cm ID port allowed evaporating liquid to X-ray tube and a tungsten-on-beryllium target, with the following escape the subcooler. The fluid was directed to an electrically settings: 200 kV, 50 milliamps, and 0.5 mm Tin filter. Images were heated “vaporizer” Vap1 which vaporized any entrained liquid and collected from 1600 pixels horizontal, 2024 pixels vertical, 0.5 s warmed the vapor to above 273 K before entering the atmo- detector exposure, averaging of 4 images per rotation position sphere. Two layers of 6.35 mm thick aerogel insulation were with a one-exposure skip and a total of 2200 rotational positions. wrapped around GV2, the hose upstream of the subcooler, the The average thickness per layer was ~15.12 µm and the subcooler itself, 3V2, 3V3, and the 3 cm length of tube between uncertainty for each layer was ± 0.7 µm. 3V3 and the subcooler to minimize heat leak into the system upstream of the test section. Instrumentation and data acquisition During the prechilling process, the liquid exiting the inner tube Next, for data acquisition (DAQ) and instrumentation, a Labview of the subcooler was directed by two “T-type” 316SS 1.27 cm ID Virtual Instrument software and National Instrument (NI) Com- three-way ball valves (3V3 and 3V2) to a fill-port on top of the pactDAQ hardware was used to collect all sensor data to be outer vessel of the subcooler. A 3 cm long, 1.270 cm OD, and displayed in real-time on a laptop. The sampling rate of all the 1.168 cm ID 304SS tube connected 3V3 to the subcooler. A sensor measurements was set to 16 Hz. Two NI-9214 TC modules pressure transducer and TC labeled “PT”, “TC” in Fig. 5a were read the signals from all the T-type TCs. NI 9205, an analog input placed between a solenoid valve (SV) and a three-way vale (3V2) module, read all the voltage signals from pressure transducers. at a distance of 7 cm from the downstream side of the inner tube The Labview VI controlled the opening and closing of the of the subcooler to measure the fluid pressure and fluid solenoidal valves (SVs), through a combination of NI-USB 6009 temperature. This station was also used to determine the and Solid-State relay. In the case of continuous flow, the relay thermodynamic state of the fluid at the inlet of the test section. energized the solenoid valve after receiving a constant voltage Once the flow inlet temperature reached a steady value, and that signal. For pulse flow, the relay energized and de-energized SV steady temperature was below the saturation temperature based according to a rectangular waveform voltage signal generated by on the measured pressure, a chilldown test was ready to the Labview VI. Signals of the two mass flow meters (Alicat M3000 commence. As shown in Fig. 5a, the test section was enclosed - SLPM) downstream of the vaporizers were read by the program in the vacuum chamber and sealed by two flanges (D and E). A directly without the NI DAQ system. 316SS vacuum chamber was used to reduce radiation and gas Fifteen TCs were soldered to the outside of each tested tube. conduction parasitic heat leak to the test section from the Five stations were spaced out axially in Fig. 5d and three TCs were surroundings, which reduces the uncertainty in the calculation of spaced out radially 90 (top, bottom, side) at each station as wall-to-fluid heat flux. A mechanical pump reduced background shown in Fig. 5e. Two cryogenic rated PTs were placed near the pressure to ~1 Pa. inlet and after the outlet of the test section by yor-lok fittings, The needle valve downstream of the test section (NV1) was respectively to provide the transient pressure histories at the two used to provide fine-tuning of the mass flow rate so that tests locations. The rest of the instrumentation is shown in Fig. 5a. could be run at different flow rates for the same dewar pressure setting. The flow was routed from the needle valve by a SS tube to Uncertainty analysis two separate vaporizers (labeled Vap2 and Vap3) that were electrically heated to vaporize the liquid-vapor two-phase flow. To Root-sum-square uncertainty analysis was conducted in a similar 27,29 enhance the heat transfer in the vaporizer, eight 1.27 cm OD fashion as in refs. ; uncertainties for test section dimensions, vacuum chamber dimensions, and thermal properties were similar copper tubes were packed inside the vaporizer in an octaweb as in . Standard error propagation rules were applied to compute configuration. One electrical heating tape was wrapped around uncertainties in chilldown time (2.1%), propellant mass consumed the vaporizer to heat it to 550 K before each test. A TC was placed at steady state (2.5%), mass flux (2.8%), and Re number (3.3%). The on the outer surface of each heating tape to monitor the temperature in real time. The flow out of Vap2 and Vap3 entered median relative uncertainties were 8–10% in Burggraf heat flux, two separate, identical gas flow meters (Gas Flow Meters 1 and 2) total heat flux, and HTC, and 25% in parasitics across all the bare that each had a capacity of 3000 standard liters per minute. The tube cases. The number of outliers in relative uncertainties were flow was then directed to the airplane vent ports downstream the on the order of 10 or fewer in each case and occurred post- flow meters. chilldown. Therefore, the 95% quantile accurately represents the maximum relative uncertainties in Burgraff heat flux, parasitics, total heat flux, and HTC which are reported in Table 1 and Test sections depicted as error bars in plots. Three, 0.914 m (36 in) long, 0.051 cm wall thickness, 1.27 cm outer diameter SS304 (properties taken from ) test sections were Experimental methodology individually flight tested: a bare tube with no coating, and a tube with a 4 layer and 7 layer coating. For the coated tubes, the SS The experimental methodology to conduct a test was as follows: tube was coated with low-thermal conductivity thin Teflon layers At the start, the needle valve was set to the target position to set on the inner surface. Specifically, the coating material was made of test section pressure, Vap 1, Vap2, and Vap3 were heated up to Fluorinated Ethylene Propylene (FEP) produced by DuPont and 550 K, and the vacuum pump system was turned on. The total classified by DuPont as Teflon 959G-203 that is a black color paint. time from engaging the pump until reaching 1 Pa inside the The coating was applied by using a pour and drain process. After vacuum chamber was ~15 min. Concurrently, the inner tube inside each pour and drain, the fresh film layer was cured in a furnace the subcooler was chilled by pressurizing the dewar, opening GV2, through a standard sintering procedure before adding another and directing the flow through 3V3 and 3V2 to the fill port of the layer by the same pour and drain procedure. As a result, the final subcooler. The subcooler took ~10 min to completely chill and fill. thickness of the coated layer depends on the total number of Then, 3V2 was shut off to stop the flow from 3V3, and the supply layers processed; for example, the 4 L coating went through the dewar was pressurized by opening the pressure regulator to the pour and drain process four separate times. To measure the desired gauge pressure for the dewar. Pressurization was done as coating layer thickness, high resolution computer tomography quickly as possible before the liquid inside the dewar could npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA J. Hartwig et al. microgravity at low Re, whether comparing 1-g pulse flow to 0-g pulse flow or 1-g coated tube flow to 1-g coated tube flow. Therefore, with optimization of coating thickness and pulse characteristics performed a priori, coated tube and pulse flow can be used for transfer line chilldown to significantly save chilldown time and mass for all future in-space cryogenic transfers. DATA AVAILABILITY Supplementary information accompanies the paper on the npj Microgravity https:// www.nature.com/npjmgrav/. Received: 16 January 2022; Accepted: 15 July 2022; Fig. 6 Performance gains of pulse flow and coated tube in microgravity: 550 kPa. Source pressure, 0.05 G level (g/g )—F2 (bare continuous) vs F3 (bare pulse 3 s 10%) vs F4 (bare pulse 2 s REFERENCES 10%) vs F13 (7 L continuous) vs F14 (7 L pulse 3 s 7%) vs F22 (4 L 1. Robbins, W. H. & Finger, H. B. An historical perspective of the NERVA nuclear continuous) vs F28 (4 L pulse 3 s 10%)—propellant mass savings as rocket engine technology program NASA-CR-187154 (1991). compared to base case of Flight_Day1_2 (continuous flow and 2. Kutter, B., Zegler, F., O’Neil, G. & Pitchford, B. A Practical, Affordable Cryogenic bare tube). Propellant Depot Based on ULA’s Flight Experience AIAA-2008-7644, 2008 SPACE Conference, 9–11 September (San Diego, CA, 2008). re-saturate at the new dewar pressure, and also before the liquid 3. Chandler, F., Bienhoff, D., Cronick, J. & Grayson, G. Propellant Depots for Earth inside the plumbing upstream of the test section could gain Orbit and Lunar Exploration AIAA-2007-6081, SPACE Conference, 18–20 Sep- enough heat to start boiling. 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V., Chandra, R., Jacob, S., Kasthurirengan, S. & Karunanithi, R. flow with a coated tube is slightly higher in microgravity (~75%) Experimental studies on cool-down and mass flow characteristics of a versus in 1-g (67%) at similar high Re. The lower mass savings in demountable liquid nitrogen transfer line. Cryogenics 36, 435–441 (1996). 1-g can easily be attributed to the fact that the duty cycle of the 19. Yuan, K., Ji, Y. & Chung, J. N. Cryogenic chilldown process under low flow rates. Int. J. Heat Mass Trans. 50, 4011–4022 (2007). 1-g coated tube pulse flow test is 20% compared to 10% of 0-g 20. Kawanami, O., Nishida, T., Honda, I., Kawashima, Y. & Ohta, H. Flow and heat coated tube pulse flow test in the current work. Lower duty cycle 33 transfer on cryogenic flow boiling during tube quenching under upward and is predicted to increase mass savings which means that a 1-g downward flow. Microgravity Sci. Tech. 19, 137–138 (2007). coated tube pulse flow test performed at 10% duty cycle would 21. Kawanami, O., Azuma, H. & Ohta, H. Effect of gravity on cryogenic boiling have >67% propellant mass savings. At high Re, the mass savings heat transfer during tube quenching. Int. J. Heat Mass Trans. 50,3490–3497 would be roughly equal in microgravity and 1-g because of forced (2007). convection dominating over buoyancy effects. However, at low Re, 22. Darr, S. R. et al. An experimental study on terrestrial cryogenic transfer line 1-g results would be expected to yield higher mass savings due to chilldown I. effect of mass flux, equilibrium quality, and inlet subcooling. Int. J. Heat Mass Trans. 103, 1225–1242 (2016). the aforementioned lack of buoyancy-assisted cooling in Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA npj Microgravity (2022) 33 J. Hartwig et al. 23. Darr, S. R. et al. An experimental study on terrestrial cryogenic transfer line ACKNOWLEDGEMENTS chilldown II. Effect of flow direction with respect to gravity and new correlation The authors acknowledge Michelle Peters and the entire flight staff and engineers at set. Int. J. Heat Mass Trans. 103, 1243–12460 (2016). the ZeroG Corporation for providing a high-quality reduced gravity environment 24. Jin, L., Park, C., Cho, H., Lee, C. & Jeong, S. Experimental investigation on chill- across multiple flights. This work was funded by the Reduced Gravity Cryogenic down process of cryogenic flow line. Cryogenics 79,96–105 (2016). Transfer Project under the Technology Demonstration Mission Program under Space 25. Jin, L., Cho, H. & Jeong, S. Experimental investigation on line chill-down process Technology Mission Directorate at NASA. by liquid argon. Cryogenics 97,31–39 (2019). 26. Jin, L., Lee, J. & Jeong, S. Investigation on heat transfer in line chill-down process with various cryogenic fluids. Int. J. Heat Mass Trans. 150, 119204 (2020). AUTHOR CONTRIBUTIONS 27. Hartwig, J. W., Styborski, J., McQuillen, J., Rame, E. & Chung, J. Liquid hydrogen J.H. contributed to developing and defining the experimental test rig, the test matrix, line chilldown experiments at high Reynolds numbers. optimal chilldown conducted data analysis, and lead author on the manuscript. The University of Florida methods. Int. J. Heat Mass Trans. 137, 703–713 (2019). team (J.N.C., J.D., B.H., H.W.) designed, built and calibrated the flight experimental 28. Hartwig, J. W., Styborski, J., Stiegemeier, B., Rame, E. & McQuillen, J. B. Liquid system. The UF team also performed both terrestrial and parabolic flight experiments hydrogen transfer line chilldown experiments. II. Analysis. Int. J. Heat Mass Trans. and collected the experimental data. J.N.C. also contributed to the development and 156, 119805 (2020). defining of the concepts of the flight experiment and assisted in the manuscript 29. Hartwig, J. W., Asensio, A. & Darr, S. R. Assessment of existing two phase heat preparation. S.D. contributed to developed and defining the test rig, lead the flight transfer coefficient and critical heat flux on cryogenic flow boiling quenching experiments, and contributed to manuscript preparation. M.T. contributed to flight experiments. Int. J. Heat Mass Trans. 93, 441–463 (2016). tests. S.J. contributed to data analysis and manuscript preparation. M.D. contributed 30. Hartwig, J. W., Hu, H., Styborski, J. & Chung, J. Comparison of cryogenic flow to developing and defining the experimental test rig. boiling in liquid nitrogen and liquid hydrogen. Int. J. Heat Mass Trans. 88, 662–673 (2015). 31. Ganesan, V., Patel, R., Hartwig, J. W. & Mudawar, I. Review of databases and COMPETING INTERESTS correlations for saturated flow boiling heat transfer coefficient for cryogens in The authors declare no competing interests. uniformly heated tubes, and development of new consolidated database and universal correlations. Int. J. Heat Mass Trans. 179, 121656 (2021). 32. Mudawar, I. Flow boiling and flow condensation in reduced gravity. Adv. Heat Trans. 49, Ch. 5 (2017). ADDITIONAL INFORMATION 33. Chung, J. N., Dong, J., Wang, H., Darr, S. R. & Hartwig, J. W. Enhancement of Supplementary information The online version contains supplementary material convective quenching heat transfer of coated tubes by intermittent cryogenic available at https://doi.org/10.1038/s41526-022-00220-9. pulse flows. Int. J. Heat Mass Trans. 141, 256–264 (2019). 34. Chung, J. N., Darr, S. R., Dong, J., Wang, H. & Hartwig, J. W. Heat transfer Correspondence and requests for materials should be addressed to Jason Hartwig. enhancement in cryogenic quenching process. Int. J. Thermal Sci. 147, 106117 (2020). Reprints and permission information is available at http://www.nature.com/ 35. Xu, W., Cheng, C. & Zhang, P. Cryogenic flow quenching of horizontal stainless reprints steel tubes. Int. J. Heat Mass Trans. (2021). 36. Chung, J. N., Dong, J., Wang, H., Darr, S. R. & Hartwig, J. W. An advance in transfer Publisher’s note Springer Nature remains neutral with regard to jurisdictional claims line chilldown heat transfer of cryogenic propellants in microgravity using in published maps and institutional affiliations. microfilm coatings for enabling deep space explorations. Nat. Microgravity 7,21 (2021). 37. Hartwig, J. W., McQuillen, J. B. & Rame, E. Pulse Chilldown Tests of a Pressure Fed Liquid Hydrogen Transfer Line AIAA-2016-2186, AIAA SciTech Conference, January 4–8 (San Diego, CA, 2016). Open Access This article is licensed under a Creative Commons 38. Shaeffer, R., Hu, H. & Chung, J. N. An experimental study on liquid nitrogen pipe Attribution 4.0 International License, which permits use, sharing, chilldown and heat transfer with pulse flows. Int. J. Heat Mass Trans. 67, 955–966 adaptation, distribution and reproduction in any medium or format, as long as you give (2013). appropriate credit to the original author(s) and the source, provide a link to the Creative 39. Burgraff, O. R. An Exact Solution of the Inverse Problem in Heat Conduction Commons license, and indicate if changes were made. The images or other third party Theory and Applications. ASME J. of Heat Trans 86, 373–380 (1964). material in this article are included in the article’s Creative Commons license, unless 40. Kirillov, P. L. et al. The Look-Up Table for Heat Transfer Coefficient in Post-Dryout indicated otherwise in a credit line to the material. If material is not included in the Region for Water Flowing in Tubes FEI-2525. (Institute of Physics and Power article’s Creative Commons license and your intended use is not permitted by statutory Engineering, Obninsk, Russia, 1996). regulation or exceeds the permitted use, you will need to obtain permission directly 41. Leung, L. K. H., Hammouda, N. & Groeneveld, D. C. A look-up table for film boiling from the copyright holder. To view a copy of this license, visit http:// th heat transfer coefficients in tubes with vertical upward flow. In Proc. 8 Inter- creativecommons.org/licenses/by/4.0/. national Topical Meeting on Nuclear Reactor Thermal-Hydraulics, 671–678 (Kyoto, Japan, 1997). 42. Marquardt, E. D., Le, J. P. & Radebaugh, R. Cryogenic Material Properties Database This is a U.S. Government work and not under copyright protection in the US; foreign 11th International Cryocooler Conference, June 20–22, (Keystone, CO, 2000). copyright protection may apply 2022 npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA http://www.deepdyve.com/assets/images/DeepDyve-Logo-lg.png npj Microgravity Springer Journals

Nitrogen flow boiling and chilldown experiments in microgravity using pulse flow and low-thermally conductive coatings

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www.nature.com/npjmgrav ARTICLE OPEN Nitrogen flow boiling and chilldown experiments in microgravity using pulse flow and low-thermally conductive coatings 1 2 2 2 2 3 3 4 Jason Hartwig , J. N. Chung , Jun Dong , Bo Han , Hao Wang , Samuel Darr , Matthew Taliaferro , Shreykumar Jain and Michael Doherty The enabling of in-space cryogenic engines and cryogenic fuel depots for future manned and robotic space exploration missions begins with technology development of advanced cryogenic fluid management systems upstream in the propellant feed system. Before single-phase liquid can flow to the engine or customer spacecraft receiver tank, the connecting transfer line must first be chilled down to cryogenic temperatures. The most direct and simplest method to quench the line is to use the cold propellant itself. When a cryogenic fluid is introduced into a warm transfer system, two-phase flow quenching ensues. While boiling is well known to be a highly efficient mode of heat transfer, previous work has shown this efficiency is lowered in reduced gravity. Due to the projected cost of launching and storing cryogens in space, it is desired to perform this chilldown process using the least amount of propellant possible, especially given the desire for reusable systems and thus multiple transfers. This paper presents an assessment of two revolutionary new performance enhancements that reduce the amount of propellant consumed during chilldown while in a microgravity environment. Twenty-eight cryogenic transfer line chilldown experiments were performed onboard four parabolic flights to examine the independent as well as combined effect of using low thermally conductive coatings and pulse flow on the chilldown process. Across a range of Reynolds numbers, results show the combination significantly enhances performance in microgravity, with a reduction in consumed mass up to 75% relative to continuous flow for a bare transfer line. npj Microgravity (2022) 8:33 ; https://doi.org/10.1038/s41526-022-00220-9 INTRODUCTION volume fill fractions in the customer receiver tank, and most in- space engines require single-phase liquid up to the injectors. The enabling of in-space cryogenic engines and cryogenic fuel Therefore, without advanced CFM technologies upstream in the depots for future manned and robotic space exploration missions feed system and storage tank, vapor ingestion is inevitable, which begins with technology development of advanced cryogenic fluid can lead to combustion instabilities within the engine. Further management (CFM) systems upstream in the propellant feed exacerbating the transfer process in microgravity is the unknown system. Cryogenic propellants offer significantly higher perfor- location of the liquid and vapor phases in the tank as well as mance relative to storable counterparts, such as hydrazine, owing reduced heat transfer. to a higher specific impulse and higher energy density. Further, Before single-phase liquid can flow to the engine or customer safety and environmental concerns over the use of toxic storable spacecraft receiver tank, the connecting transfer line must first be propellants have led to the ongoing examination of more “green” chilled down to cryogenic temperatures. Chilldown, or quenching, propellants such as liquid methane as alternate fuel sources. Aside is defined as the transient process of cooling hardware down to from nuclear thermal propulsion systems , no other known pure cryogenic temperatures so that vapor-free liquid can eventually chemical propulsion system propellant combination can deliver a flow between two points of interest. The most direct and simplest higher ISP than liquid hydrogen/liquid oxygen. However, there are method to quench the line is to use the cold propellant itself. challenging aspects when working with cryogens due to inherent When a cryogenic fluid is initially transferred through a system, thermo-physical properties. Particularly for the current work, the the tube walls and hardware (e.g. valves) undergo a transient low normal boiling point (NBP), low surface tension, and high chilldown prior to reaching a steady state of operation. Chilldown susceptibility to parasitic heat leak leads to unwanted boiling and thus involves unsteady two-phase heat and mass transfer and flow two-phase flow during propellant transfer. boiling. While boiling is well known to be a highly efficient mode 2,3 Cryogenic fuel depots ,defined as an Earth-orbiting propellant of heat transfer, previous work has shown this efficiency is storage vessel that would house cryogenic propellant to allow significantly lowered in reduced gravity, both for room tempera- 4–8 spacecraft to refuel, have four stages: (1) acquisition of the storage ture fluids as well as cryogens . Due to the projected cost of tank liquid, (2) chilldown of the connecting transfer line hardware, launching and storing cryogens in space, it is desired to perform (3) chilldown of the receiver tank, and (4) fill of the receiver tank, this chilldown process using the least amount of propellant as all in the microgravity of space; this paper focuses on the second possible, especially given the drive towards reusable systems and stage, chilldown of the transfer line. Meanwhile, cryogenic engines thus multiple transfers. also require acquisition of the storage tank liquid and chilldown of Numerous cryogenic flow boiling quenching experiments have the transfer line. Cryogenic fuel depots will require very high liquid previously been conducted on bare tubes, the results up to 2018 1 2 3 4 NASA Glenn Research Center, Cleveland, OH 44135, USA. University of Florida, Gainesville, FL 32611, USA. The Aerospace Corporation, El Segundo, CA 90245, USA. Georgia Tech University, Atlanta, GA 30332, USA. email: Jason.W.Hartwig@nasa.gov Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA 1234567890():,; J. Hartwig et al. of which are summarized in ref. . Key contributions in 1-g were conductive coatings and pulse flow on the chilldown process. 4,5,10–23 provided by , which investigated the effect of mass flux, While previous experiments have reported the effects of pulse inlet state, pressure, and flow direction on cryogenic tube flow and coatings on transfer line chilldown in Earth-gravity, this is chilldown, predominately using liquid nitrogen (LN ) and liquid the first report of pulse flow and the combined effect of pulse flow hydrogen (LH ). Since 2018, five more cryogenic quenching with a coated tube in a microgravity environment. studies on bare tubes that passed the data filtering criteria from ref. by Jin et al. have been added to the cryogenic database to 24 25 RESULTS AND DISCUSSION cover low Reynolds (Re) number LN , liquid argon , and liquid 26 27,28 oxygen chilldown experiments while added high Re number Test matrix chilldown tests with LH . Hartwig et al. recently summarized Table 1 lists the complete flight test matrix. Ground tests were cryogenic quenching flow boiling trends over the consolidated performed at the University of Florida, while flight tests were literature, across multiple flow regimes, mass fluxes, inlet states, conducted during the low-gravity portion of the classic parabolic and gravity levels. For all cryogens, the chilldown process is highly trajectory followed by the flight provider ZeroG. Pressure is the dominated by the film boiling regime for bare tubes (for measured pressure at the inlet to the test section, time-averaged quantum fluids such as hydrogen and helium, there are additional over the test duration. Period is the sum of valve “on” and “off” time factors at play). When a cryogen is introduced into a warm tube, for a pulse flow test cycle. The duty cycle is the ratio of the valve “on” especially at high mass flux and low inlet equilibrium quality, a time to the period. For example, with a period of 3 s and a duty cycle vapor film blanket surrounds the liquid core which acts as an of 10%, the valve is on 0.3 s and off for 2.7 s. G level is the gravity insulator that inhibits heat transfer between cold liquid and warm level as read by accelerometers attached to the experimental rig tube. At lower mass flux and saturated inlet states, dryout occurs while on the flight. Note that a few of the flight tests were conducted over a longer distance along the tube as in the case of traditional at a g-level higher than nominal; these were deemed “Martian gravity fluids . Film boiling heat transfer is a highly inefficient process tests”. Coating thickness in number of layers, “L”, is described in the relative to transition and nucleate boiling. In most instances, film “Methods” section. boiling can persist for >85% of the total time needed to chill the Chilldown time was determined as follows: In practice, the most tube down to the saturation temperature of the cryogen. Once the stringent chilldown criteria would be determined from a Leidenfrost point is reached, chilldown proceeds into transition measured stream temperature downstream of the test section boiling, nucleate boiling, and then single-phase liquid convective reading lower than the saturation temperature based on the flow. In microgravity, this poor heat transfer is exacerbated by the downstream pressure; however, this measurement was not lack of buoyancy force; cryogenic film boiling heat transfer was available for the current tests. Based on boiling heat transfer shown to be 25% lower at low to modest Re flows relative to 1-g . theory, nucleate boiling would end when the inner surface At very high Re, inertial forces can overcome gravitational forces temperature drops below that of the onset-of-nucleate boiling such that gravity no longer affects flow boiling (although this (ONB). As a result, the wall heat flux would switch from higher has not been demonstrated yet for cryogens). boiling heat flux to much lower single-phase convective heat flux To overcome this hurdle in poor performance, researchers have that would reflect a change on the outer wall surface temperature recently investigated low thermally conductive materials applied gradient with time. A computed inner wall temperature could also to the inner tube walls and the effect of such coatings on the not be used to determine end of chilldown; while the inverse chilldown process. The coating acts as an insulator between the conduction method of Burggraf can be used to determine inner cold propellant and warm wall, resulting in an inner wall surface wall temperature for bare tubes, due to the unknown thermal temperature that reaches the Leidenfrost point without cooling contact resistances between the coated layer and tube inner wall the entire tube mass. Recent 1-g experiments conducted in the 33,34 35 as well as among adjacent coated layers, inner wall temperature United States and China independently confirmed that a could not be determined for coated tubes. Therefore, outer wall Teflon coated tube could reduce chilldown times up to 75% over temperature data had to be used to determine end of chilldown. an uncoated stainless steel (SS) tube using LN . Both researchers Three chilldown criteria were explored: (1) the averaged exit outer also investigated the effect of the Teflon coating thickness on wall temperature was compared to the liquid saturation chilldown performance, and both showed that thicker coatings temperature (based on local downstream pressure), (2) the first led to faster chilldown times. However, as the coating thickness derivative of outer wall temperature (with respect to time) increased further there was an apparent point of diminishing reaching and remaining near 0 K/s (due to minimal convective returns because the chilldown curves (wall temperature versus heat transfer between single-phase liquid and tube), and (3) a time) converged at the highest tested thicknesses. Coated tubes peak value in the second derivative of outer wall temperature offer hope to combat the intrinsically poor film boiling heat (with respect to time) which would indicate the slope change in transfer in microgravity . the chilldown curve occurring at onset of nucleate boiling (ONB). A second way to enhance poor chilldown performance is to use 37 38 The first method was found to be unreliable due to inaccurate pulse flow. Demonstrated using both LH and LN , in pulse 2 2 chilldown time estimations attributed to the significant difference flow, the inlet valve is cyclically opened and closed with a between inner and outer wall temperature at higher layers of specified duty cycle (DC) and pulse width until the desired degree coating. The third method also yielded inaccurate chilldown time of chilldown is reached. The advantage of pulse flow is lower mass estimations attributed to the absence of a true global maxima in consumption over traditional continuous flow due to more the second derivative at higher layers of coating. Therefore, the efficient usage of latent and sensible energy of the fluid, with second method using the first derivative (typically using test the disadvantage being potential valve fatigue and/or failure and section averaged temperature), slightly conservative but consis- added complexity in operation. tent across all scenarios, was used to determine the end of The purpose of this paper is to present an assessment of two chilldown in all test cases. new performance enhancements that reduce the amount of Chilldown mass was the total consumed LN mass at the end of propellant consumed during chilldown while in a microgravity 2 chilldown as read by the flow meter downstream of the test environment, and to investigate if the mass savings holds in section: microgravity. Twenty-eight LN transfer line chilldown experi- ments were performed onboard a parabolic flight that simulated Z end space microgravity conditions to examine the independent as well m ¼ mdt _ (1) LN2 as combined performance gains of using low thermally 0 npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA 1234567890():,; J. Hartwig et al. Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA npj Microgravity (2022) 33 Table 1. Flight test matrix. # Test name Inlet pressure Period [s] Duty G Level [g/g ] Coating Chilldown Total chilldown Steady-state Re Steady-state mass cycle [%] thickness time [s] mass [kg] flux [kg/m -s] F1 Flight_day1_1 550 kPa [80 psia] – 100 0.38 Bare 15.1 1.08 71,976 548.87 F2 Flight_day1_2 550 kPa [80 psia] – 100 0.05 Bare 14.6 0.872 67,324 533.22 F3 Flight_day1_3 550 kPa [80 psia] 2 10 0.05 Bare 16 0.597 53,859 426.58 F4 Flight_day1_4 550 kPa [80 psia] 3 10 0.05 Bare 16.7 0.614 64,092 507.63 F5 Flight_day1_5 340 kPa [50 psia] – 100 0.05 Bare 70.6 0.264 1761 22.97 F6 Flight_day1_6 340 kPa [50 psia] 4 10 0.05 Bare 78.1 0.279 1383 18.04 F7 Flight_day2_1 550 kPa [80 psia] 3 7 0.38 Bare 21.3 0.637 32,621 347.47 F8 Flight_day2_2 340 kPa [50 psia] – 100 0.05 Bare 80.3 0.279 1566 20.42 F9 Flight_day2_3 340 kPa [50 psia] 5 4 0.05 Bare 162 0.304 881 11.49 F10 Flight_day2_4 340 kPa [50 psia] – 100 0.05 Bare 172 0.16 992 15.94 F11 Flight_day2_5 440 kPa [65 psia] 3 7 0.05 Bare 21.3 0.645 29,761 306.99 F12 Flight_day2_6 440 kPa [65 psia] – 100 0.05 Bare 21.1 1.174 53,021 459.65 F13 Flight_day3_1 550 kPa [80 psia] – 100 0.05 7 L 8.19 0.469 63,912 506.19 F14 Flight_day3_2 550 kPa [80 psia] 3 7 0.05 7 L 9.94 0.339 38,308 351.49 F15 Flight_day3_3 550 kPa [80 psia] – 100 0.05 7 L 6.69 0.309 45,760 340.84 F16 Flight_day3_4 550 kPa [80 psia] 3 7 0.05 7 L 10 0.248 43,178 329.21 F17 Flight_day3_5 340 kPa [50 psia] – 100 0.05 7 L 12.8 0.0881 8945 73.36 F18 Flight_day3_6 440 kPa [65 psia] 3 7 0.05 7 L 9.31 0.283 38,265 364.8 F19 Flight_day3_7 440 kPa [65 psia] – 100 0.05 7 L 7.06 0.514 83,353 649.18 F20 Flight_day3_8 440 kPa [65 psia] – 100 0.05 7 L 7.88 0.585 88,127 682.68 F21 Flight_day3_9 340 kPa [50 psia] – 100 0.05 7 L 12.8 0.0855 10,949 90.82 F22 Flight_day4_1 550 kPa [80 psia] – 100 0.05 4 L 4.69 0.281 74,196 558.16 F23 Flight_day4_2 550 kPa [80 psia] 3 7 0.05 4 L 9.06 0.269 46,373 348.85 F24 Flight_day4_3 550 kPa [80 psia] 2 10 0.05 4 L 6.19 0.237 63,067 474.44 F25 Flight_day4_4 440 kPa [65 psia] – 100 0.05 4 L 4.5 0.23 59,931 476.14 F26 Flight_day4_5 440 kPa [65 psia] 3 7 0.05 4 L 10 0.288 50,941 404.72 F27 Flight_day4_6 340 kPa [50 psia] – 100 0.05 4 L 7.81 0.0555 8061 66.22 F28 Flight_day4_7 550 kPa [80 psia] 3 10 0.05 4 L 7.25 0.274 64,048 500.43 These cases were conducted over multi-parabolas; the calculations also include the high G part. Airplane accelerated to 1 g before end of chilldown. The initial temperature is lower than other cases because there was not enough time to reheat. Missing time-series mass flow rate data. J. Hartwig et al. Fig. 1 Test F2 (Flight_Day1_2, continuous flow, 550 kPa source pressure, 67324 Re, bare tube surface, 0.05 G level (g/g0)). a Chilldown curve based on average exit wall temperature, b Boiling curve based on average exit wall temperature, c Chilldown curve of all TCs. where t is the end of chilldown time and mðtÞ is the time- Averaging was done by adding the temperatures and dividing by end dependent LN mass flow rate measured by the gas flow meter. the number of sensors. Figure 1c illustrates the chilldown curve of Steady state Reynolds (Re) number (defined at the end of all thermocouples (TCs) placed on the tube outer wall according to chilldown when single phase liquid flow is established) and mass Fig. 5d, e in the “Methods” section. Errors bars are plotted but flux were evaluated using inner diameter and saturation condi- barely discernable. Three boiling regimes, film boiling (FB), tions based on the measured test section pressure: transition boiling (TB), and nucleate boiling (NB), and single- phase convection are separated by three critical points, the 4m _ Re ¼ (2) Leidenfrost Point (LFP), Critical Heat Flux (CHF), and the onset of πDμ nucleate boiling (ONB). The chilldown curve begins in the film 39 boiling regime where the cold liquid entering the warm tube For uncoated tubes, the method of Burgraff was used to experiences violent boiling. Depending on the local conditions, determine inner wall temperature and transient radial heat the flow will proceed into dispersed flow FB (high quality, low conduction through the tube as follows: subcooling, low mass flux) or inverted annular FB (low quality, 2 2 2 3 4 2 r r ðÞ ρc r r r r 00 o dT P i r d T 40,41 i o i o o i o high subcooling, high mass flux) . The high wall surface q ¼ ρc þ   ln P 2 2r dt k 16 16r 4 r dt i i o temperature causes the liquid to completely vaporize before 3 5 4 2 3 6 2 3 4 r 3r r r r ðÞ ρc 3r r r r r d T P i o i o i o o i ri o i ri o þ  þ   ln  ln reaching the surface resulting in an inner liquid core and outer 2 3 k 384 128 128 384r 128 r 32 r dt i o o annular vapor core. This vapor blanket along the wall insulates the (3) warm pipe from the cold liquid, causing the temperature of the where q is the radial heat flux through the tube, ρ, C , and k are pipe to decrease, albeit slowly. FB is the least efficient quenching the tube density, specific heat, and thermal conductivity, mechanism. As the transfer line chills down, the system respectively, r and r are the inner and outer radii, and T is the approaches the LFP, or rewet temperature, where heat flux is at i o o outer wall temperature. Heat transfer coefficient was then a minimum (during boiling). Heat transfer here is a minimum due computed as follows: to the inefficiency of heat transfer between cold vapor and wall. LFP is also characterized by the onset of a rapid drop in wall 00 00 o 00 00 00 q þ q þ q þ q þ q axial r rad solidcond gascond temperature. As shown in Fig. 1c, the LFP occurs at later times for (4) h ¼ quench TCs located farther downstream. This trend demonstrates the T  T i sat location of the quenching front as it propagates downstream as where q is the axial conduction along the tube; the terms in axial chilldown evolves. The flow then proceeds and passes quickly parenthesis are the radiation, solid conduction, and gaseous through TB, characterized by intermittent liquid contact along the conduction parasitic heat leak terms, respectively, T is the inner walls. TB ends when liquid is in full contact with the walls at the wall temperature which comes from Burgraff’s method, and T is sat point of CHF. Heat transfer is a maximum at CHF due to the highly the saturation temperature based on the measured pressure. The efficient cooling process of boiling. Nucleate boiling follows, method to calculate the different heat fluxes in Eq. 4 has been where heat is transferred by vapor bubbles formed in surface 22 27 shown in many other papers, see for example ref. and ref. . cavities that are swept away from the tube surface. Depending on the inlet conditions, NB can be liquid-convection dominate or Governing physics of chilldown nucleation-dominate . As the wall cools further, the tube inner Figure 1a shows the chilldown curve of averaged exit wall surface approaches the ONB, characterized as the point at which temperature (TC5, TC10, TC15) and Fig. 1b shows the boiling curve the system evolves from nucleate two-phase cooling to single- based on the averaged exit wall temperature in microgravity. phase liquid convection and an obvious slope change in the npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA J. Hartwig et al. Fig. 2 Effect of teflon coating thickness in microgravity: 550 kPa source pressure, continuous flow, 0.05 G level (g/g0). F2 (Flight_Day1_2, 67324 Re, bare surface) versus F22 (Flight_Day4_1, 74196 Re, 4 L coating) versus F13 (Flight_Day3_1, 63912 Re, 7 L coating) a Chilldown curve, b exit pressure, c mass flux, and d total propellant mass consumed based on averaged exit wall temperature. chilldown curve. Vapor-free liquid marks the end of the chilldown wall temperature for the coated tubes will be significantly lower test. The single-phase cooling causes the wall temperature to drop since the coating restricts the heat transfer between inner and slowly to the liquid saturation temperature and then remain outer walls. Second, less mass is consumed for coated over bare steady as heat transfer reduces to near zero. In microgravity, tubes as substantiated in Fig. 2d; the 4 L and 7 L coated cases have circumferential TCs at each station have almost identical chilldown 68 and 46% propellant mass savings over the bare tube, behavior at any axial distance from the inlet; stratification effects respectively. normally seen for horizontal tubes in 1-g disappear, leading to Third, however, there is an apparent point of diminishing axisymmetric flow patterns through the tube, and thus uniform returns; this trend of improved chilldown performance upon chilldown circumferentially. addition of coating is reversed when the number of coating layers is increased from 4 to 7 because the chilldown time is faster for 4 L (4.7 s) compared to 7 L (8.2 s) case. Similarly, from 4 L to 7 L, the Bare vs. coated tube in microgravity, continuous flow propellant mass savings and chilldown efficiency are reduced. This Figure 2a–d plot chilldown curves, exit pressure, mass flux, and crossover in performance and possible existence of an optimal total consumed liquid mass for bare, 4 L, and 7 L coated tubes for coating layer is explained by counteracting heat transfer higher steady state Re (63,912–74,196). The initial fluctuations in mechanisms: (1) the low thermal conductivity of the coating layer pressure measurements in Fig. 2b are due to the transient nature facilitates the faster temperature drop of tube inner surface by of the flow at start of the test. Shortly after the transient start, restricting heat transfer between inner surface and bulk of the downstream pressure measurements reach their steady-state metal tube and (2) the low thermal conductivity coating also value and remain there until at least the end of chilldown in all creates a thermal resistance that restricts the heat conduction three cases. The mass flux of the 4 L coating case in Fig. 2cisa between bulk of the tube and cooling fluid. With these contrasting straight horizontal line because of missing timed mass flow rate mechanisms at play, the thickness of the coating must be such data for that run; a linear correlation was developed between that it is thick enough to quickly lower the tube inner surface averaged inlet pressure and averaged mass flow rate for cases temperature while being thin enough to facilitate fast wall with available mass flow rate data that were run at 0.05 g level and chilldown. However, the presence of the coating accelerates were completed under one parabola. This linear correlation was chilldown as evident in any comparison between bare and coated then used to calculate an average mass flow rate for cases with tube at similar thermodynamic conditions. missing mass flow rate data (but available inlet pressure data). Trends are as follows: First, coating the inner wall of the tube Continuous versus pulse flow in microgravity, bare tube drastically affects the chilldown behavior and leads to faster chilldown times. The low thermally conductive Teflon layer acts as Figure 3a–e plot chilldown curves, heat transfer coefficient, an insulator between cold fluid and warm wall; the inner surface pressure, mass flux, and total consumed liquid mass for temperature chills down quickly without cooling the entire tube continuous flow and pulsed flow at a period of 2 s and duty mass. The lower inner wall surface temperature earlier on means cycle 10% (valve on 0.2 s, valve off 1.8 s) and for period 3 s and that the Leidenfrost point is reached faster such that the liquid can duty cycle 10% (valve on 0.3 s, valve off 2.7 s) at higher Re stay in contact with the tube for the heat transfer to be in TB and (53859–64092). Trends are as follows: First, both continuous and NB that reduces the poor heat transfer film boiling time; this is pulse flow exhibit the same chilldown curve and proceed through substantiated by the drastic slope change for 4 L and 7 L tube the same transition points. For pulse flow, the longer the valve-off indicating the LFP is reached earlier on relative to the bare tube. time, the more the tube temperature stabilizes as residual cooling Note that Fig. 2a plots outer wall temperature; the actual inner due to blowdown diminishes. Second, from Fig. 3a, e, it is clear Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA npj Microgravity (2022) 33 J. Hartwig et al. Fig. 3 Effect of pulse flow on a bare tube in microgravity: 550 kPa source pressure, bare tube surface, 0.05 G level (g/g0)—F2 (Flight_Day1_2, 67324 Re, continuous) versus F3 (Flight_Day1_3, 53859 Re, pulse 2 s 10%) versus F4 (Flight_Day1_4, 64092 Re, pulse 3 s 10%). a Chilldown curve, b Heat transfer coefficient versus wall superheat, c Exit pressure, d mass flux, and e total propellant mass consumed based on averaged exit wall temperature. Fig. 4 Combined effect of coatings and pulse flow in microgravity: 0.05 G level (g/g0)—F2 (Flight_Day1_2, 550 kPa source pressure, 67,324 Re, bare surface, continuous) versus F28 (Flight_Day4_7, 550 kPa source pressure, 64,048 Re, 4 L Coating, 3 s 10%). a Chilldown curve and b total propellant mass consumed based on averaged exit wall temperature. that pulse flow achieves chilldown using less propellant but at the Table 1, there is 29–32% mass savings with pulse flow in cost of longer chilldown time due to better use of sensible and comparison to continuous flow at these flight conditions. Third, latent energy of the fluid. Figure 3c, d shows fluctuations in for a fixed duty cycle, reducing the valve-open time leads to pressure and mass flux that are due to valve cycling, that these slightly shorter chilldown times (although not shown directly in fluctuations continue until end of chilldown, and that the Fig. 3) and, slightly less propellant consumption as shown in fluctuation amplitudes are higher for longer periods. From Fig. 3a, e; this trend compares well with previous pulse flow tests npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA J. Hartwig et al. Fig. 5 Experimental Design. a Flight system piping and instrumentation diagram, b actual flight line chilldown rig, c pre-cooler, d test section, vacuum chamber, and thermocouple locations, A: Inlet, B: 3.81 cm (1.5 in) long tube section, C: ultra-torr fitting, D: left flange, E: right flange and e thermocouple locations at each station. 33 37 for both LN and LH . Fourth, for this particular comparison, METHODS 2 2 Fig. 3b shows that continuous flow exhibited a higher CHF over Experimental description pulse flow, and that reducing the valve open time reduced the The authors have completed four successful cryogenic line CHF. Because of the temperature stabilization when the valve was chilldown parabolic flight campaigns including the current cycled off, the temperature does not drop as rapidly in pulse campaign between 2015 and 2020 and are familiar with system compared to continuous flow which caused the wall temperature designs, troubleshooting, issues, and failures that arise with first derivative term to be lower at CHF in pulse flow. However, if microgravity flight testing. The fourth-generation system was 8 36 the CHF was traversed when the valve was on, it is expected that modified based on flights from the first- and second-generation the pulse flow heat transfer coefficient (HTC) would be nearly systems. As before, the system is intended for both ground and equivalent to that of continuous flow. For bare tubes in flight experiments. Figure 5a shows a system flow network and piping and instrumentation diagram while Fig. 5b shows a picture microgravity, while higher frequency, shorter pulse widths are of the actual flight rig. favorable from a chilldown efficiency standpoint, more valve LN was supplied to the system from an 80-liter vacuum- cycles implies higher risk of valve degradation and potential jacketed dewar, with a relief valve set at 861 kPa. A gaseous failure. Therefore, there is an inherent trade-off in which the nitrogen (GN ) cylinder initially pressurized at 15 MPa was used to optimal valve duty cycle could be determined. pressurize the dewar to a set value for each test, which ranged between 90 and 830 kPa absolute pressure. Dewar pressure was managed by a pressure regulator that controlled the dewar Performance gain of combined pulsed flow and coated tubes pressure to within 35 kPa of the set value during each test. in microgravity Depressurization was carried out by opening the globe valve 2 Figure 4a, b plot chilldown curves and total consumed liquid mass (GV2) and the three-way ball valve 1 (3V1) to allow ullage gas to for bare tube with continuous flow and 4 L coated tube with vent to the atmosphere. pulsed flow characterized by a period of 3 sec and duty cycle of The dewar was used to supply the LN both for prechilling the 10% (valve on 0.3 s, valve off 2.7 s) at higher Re (64,048–67,324). plumbing upstream of the test section and for conducting the actual chilldown experiment. LN was delivered through valve Trends are as follows: First, the effect of coating on reducing GV3 that was connected through a 1.2 m long, 1.27 cm outer chilldown time seems to outweigh the effect of pulse flow on diameter (OD), 1.18 cm inner diameter (ID) 304 (stainless steel) SS increasing chilldown time as evidenced by the sharp drop in braided hose to the inner tube of the precooler (or subcooler) temperature at ~4 s in Fig. 4a for the coated tube. Second, the shell-tube heat exchanger shown in Fig. 5c. The subcooler served individual benefits of propellant mass savings with coating and three purposes: (1) to preserve subcooling of the liquid from the pulse flow are nearly perfectly superimposed, leading to a 76% storage tank flowing through the transfer line by eliminating reduction in propellant consumption. Results thus show that high parasitic heat leak, (2) to slightly subcool the LN in the transfer performance is still achieved in microgravity for pulse flow with a line, since the saturation temperature of the shell side was always low thermal conductivity coating which leads to a reduction in lower than the tube side, and most importantly (3) to ensure chilldown time and mass and increase in chilldown efficiency over single-phase liquid at the inlet of the test section. The liquid level continuous flow with a bare tube. of the nitrogen pool was monitored by three thermocouples (TC) Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA npj Microgravity (2022) 33 J. Hartwig et al. inside the subcooler, two at the shell-side and one at the outlet to x-ray scans of the tube cross sections were obtained using a Vap1. The temperature readings of these TCs were displayed on a Phoenix v|tome|x M system in the Nano Research Facility at the laptop in real-time. The level of the LN pool was inferred from the University of Florida. Scanning was carried out using a 240 kV TC insertion depth. A 2.5 cm ID port allowed evaporating liquid to X-ray tube and a tungsten-on-beryllium target, with the following escape the subcooler. The fluid was directed to an electrically settings: 200 kV, 50 milliamps, and 0.5 mm Tin filter. Images were heated “vaporizer” Vap1 which vaporized any entrained liquid and collected from 1600 pixels horizontal, 2024 pixels vertical, 0.5 s warmed the vapor to above 273 K before entering the atmo- detector exposure, averaging of 4 images per rotation position sphere. Two layers of 6.35 mm thick aerogel insulation were with a one-exposure skip and a total of 2200 rotational positions. wrapped around GV2, the hose upstream of the subcooler, the The average thickness per layer was ~15.12 µm and the subcooler itself, 3V2, 3V3, and the 3 cm length of tube between uncertainty for each layer was ± 0.7 µm. 3V3 and the subcooler to minimize heat leak into the system upstream of the test section. Instrumentation and data acquisition During the prechilling process, the liquid exiting the inner tube Next, for data acquisition (DAQ) and instrumentation, a Labview of the subcooler was directed by two “T-type” 316SS 1.27 cm ID Virtual Instrument software and National Instrument (NI) Com- three-way ball valves (3V3 and 3V2) to a fill-port on top of the pactDAQ hardware was used to collect all sensor data to be outer vessel of the subcooler. A 3 cm long, 1.270 cm OD, and displayed in real-time on a laptop. The sampling rate of all the 1.168 cm ID 304SS tube connected 3V3 to the subcooler. A sensor measurements was set to 16 Hz. Two NI-9214 TC modules pressure transducer and TC labeled “PT”, “TC” in Fig. 5a were read the signals from all the T-type TCs. NI 9205, an analog input placed between a solenoid valve (SV) and a three-way vale (3V2) module, read all the voltage signals from pressure transducers. at a distance of 7 cm from the downstream side of the inner tube The Labview VI controlled the opening and closing of the of the subcooler to measure the fluid pressure and fluid solenoidal valves (SVs), through a combination of NI-USB 6009 temperature. This station was also used to determine the and Solid-State relay. In the case of continuous flow, the relay thermodynamic state of the fluid at the inlet of the test section. energized the solenoid valve after receiving a constant voltage Once the flow inlet temperature reached a steady value, and that signal. For pulse flow, the relay energized and de-energized SV steady temperature was below the saturation temperature based according to a rectangular waveform voltage signal generated by on the measured pressure, a chilldown test was ready to the Labview VI. Signals of the two mass flow meters (Alicat M3000 commence. As shown in Fig. 5a, the test section was enclosed - SLPM) downstream of the vaporizers were read by the program in the vacuum chamber and sealed by two flanges (D and E). A directly without the NI DAQ system. 316SS vacuum chamber was used to reduce radiation and gas Fifteen TCs were soldered to the outside of each tested tube. conduction parasitic heat leak to the test section from the Five stations were spaced out axially in Fig. 5d and three TCs were surroundings, which reduces the uncertainty in the calculation of spaced out radially 90 (top, bottom, side) at each station as wall-to-fluid heat flux. A mechanical pump reduced background shown in Fig. 5e. Two cryogenic rated PTs were placed near the pressure to ~1 Pa. inlet and after the outlet of the test section by yor-lok fittings, The needle valve downstream of the test section (NV1) was respectively to provide the transient pressure histories at the two used to provide fine-tuning of the mass flow rate so that tests locations. The rest of the instrumentation is shown in Fig. 5a. could be run at different flow rates for the same dewar pressure setting. The flow was routed from the needle valve by a SS tube to Uncertainty analysis two separate vaporizers (labeled Vap2 and Vap3) that were electrically heated to vaporize the liquid-vapor two-phase flow. To Root-sum-square uncertainty analysis was conducted in a similar 27,29 enhance the heat transfer in the vaporizer, eight 1.27 cm OD fashion as in refs. ; uncertainties for test section dimensions, vacuum chamber dimensions, and thermal properties were similar copper tubes were packed inside the vaporizer in an octaweb as in . Standard error propagation rules were applied to compute configuration. One electrical heating tape was wrapped around uncertainties in chilldown time (2.1%), propellant mass consumed the vaporizer to heat it to 550 K before each test. A TC was placed at steady state (2.5%), mass flux (2.8%), and Re number (3.3%). The on the outer surface of each heating tape to monitor the temperature in real time. The flow out of Vap2 and Vap3 entered median relative uncertainties were 8–10% in Burggraf heat flux, two separate, identical gas flow meters (Gas Flow Meters 1 and 2) total heat flux, and HTC, and 25% in parasitics across all the bare that each had a capacity of 3000 standard liters per minute. The tube cases. The number of outliers in relative uncertainties were flow was then directed to the airplane vent ports downstream the on the order of 10 or fewer in each case and occurred post- flow meters. chilldown. Therefore, the 95% quantile accurately represents the maximum relative uncertainties in Burgraff heat flux, parasitics, total heat flux, and HTC which are reported in Table 1 and Test sections depicted as error bars in plots. Three, 0.914 m (36 in) long, 0.051 cm wall thickness, 1.27 cm outer diameter SS304 (properties taken from ) test sections were Experimental methodology individually flight tested: a bare tube with no coating, and a tube with a 4 layer and 7 layer coating. For the coated tubes, the SS The experimental methodology to conduct a test was as follows: tube was coated with low-thermal conductivity thin Teflon layers At the start, the needle valve was set to the target position to set on the inner surface. Specifically, the coating material was made of test section pressure, Vap 1, Vap2, and Vap3 were heated up to Fluorinated Ethylene Propylene (FEP) produced by DuPont and 550 K, and the vacuum pump system was turned on. The total classified by DuPont as Teflon 959G-203 that is a black color paint. time from engaging the pump until reaching 1 Pa inside the The coating was applied by using a pour and drain process. After vacuum chamber was ~15 min. Concurrently, the inner tube inside each pour and drain, the fresh film layer was cured in a furnace the subcooler was chilled by pressurizing the dewar, opening GV2, through a standard sintering procedure before adding another and directing the flow through 3V3 and 3V2 to the fill port of the layer by the same pour and drain procedure. As a result, the final subcooler. The subcooler took ~10 min to completely chill and fill. thickness of the coated layer depends on the total number of Then, 3V2 was shut off to stop the flow from 3V3, and the supply layers processed; for example, the 4 L coating went through the dewar was pressurized by opening the pressure regulator to the pour and drain process four separate times. To measure the desired gauge pressure for the dewar. Pressurization was done as coating layer thickness, high resolution computer tomography quickly as possible before the liquid inside the dewar could npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA J. Hartwig et al. microgravity at low Re, whether comparing 1-g pulse flow to 0-g pulse flow or 1-g coated tube flow to 1-g coated tube flow. Therefore, with optimization of coating thickness and pulse characteristics performed a priori, coated tube and pulse flow can be used for transfer line chilldown to significantly save chilldown time and mass for all future in-space cryogenic transfers. DATA AVAILABILITY Supplementary information accompanies the paper on the npj Microgravity https:// www.nature.com/npjmgrav/. Received: 16 January 2022; Accepted: 15 July 2022; Fig. 6 Performance gains of pulse flow and coated tube in microgravity: 550 kPa. 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Tech. 19, 137–138 (2007). coated tube pulse flow test performed at 10% duty cycle would 21. Kawanami, O., Azuma, H. & Ohta, H. Effect of gravity on cryogenic boiling have >67% propellant mass savings. At high Re, the mass savings heat transfer during tube quenching. Int. J. Heat Mass Trans. 50,3490–3497 would be roughly equal in microgravity and 1-g because of forced (2007). convection dominating over buoyancy effects. However, at low Re, 22. Darr, S. R. et al. An experimental study on terrestrial cryogenic transfer line 1-g results would be expected to yield higher mass savings due to chilldown I. effect of mass flux, equilibrium quality, and inlet subcooling. Int. J. Heat Mass Trans. 103, 1225–1242 (2016). the aforementioned lack of buoyancy-assisted cooling in Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA npj Microgravity (2022) 33 J. Hartwig et al. 23. Darr, S. R. et al. An experimental study on terrestrial cryogenic transfer line ACKNOWLEDGEMENTS chilldown II. Effect of flow direction with respect to gravity and new correlation The authors acknowledge Michelle Peters and the entire flight staff and engineers at set. Int. J. Heat Mass Trans. 103, 1243–12460 (2016). the ZeroG Corporation for providing a high-quality reduced gravity environment 24. Jin, L., Park, C., Cho, H., Lee, C. & Jeong, S. Experimental investigation on chill- across multiple flights. This work was funded by the Reduced Gravity Cryogenic down process of cryogenic flow line. Cryogenics 79,96–105 (2016). Transfer Project under the Technology Demonstration Mission Program under Space 25. Jin, L., Cho, H. & Jeong, S. Experimental investigation on line chill-down process Technology Mission Directorate at NASA. by liquid argon. Cryogenics 97,31–39 (2019). 26. Jin, L., Lee, J. & Jeong, S. Investigation on heat transfer in line chill-down process with various cryogenic fluids. Int. J. Heat Mass Trans. 150, 119204 (2020). AUTHOR CONTRIBUTIONS 27. Hartwig, J. W., Styborski, J., McQuillen, J., Rame, E. & Chung, J. Liquid hydrogen J.H. contributed to developing and defining the experimental test rig, the test matrix, line chilldown experiments at high Reynolds numbers. optimal chilldown conducted data analysis, and lead author on the manuscript. The University of Florida methods. Int. J. Heat Mass Trans. 137, 703–713 (2019). team (J.N.C., J.D., B.H., H.W.) designed, built and calibrated the flight experimental 28. Hartwig, J. W., Styborski, J., Stiegemeier, B., Rame, E. & McQuillen, J. B. Liquid system. The UF team also performed both terrestrial and parabolic flight experiments hydrogen transfer line chilldown experiments. II. Analysis. Int. J. Heat Mass Trans. and collected the experimental data. J.N.C. also contributed to the development and 156, 119805 (2020). defining of the concepts of the flight experiment and assisted in the manuscript 29. Hartwig, J. W., Asensio, A. & Darr, S. R. Assessment of existing two phase heat preparation. S.D. contributed to developed and defining the test rig, lead the flight transfer coefficient and critical heat flux on cryogenic flow boiling quenching experiments, and contributed to manuscript preparation. M.T. contributed to flight experiments. Int. J. Heat Mass Trans. 93, 441–463 (2016). tests. S.J. contributed to data analysis and manuscript preparation. M.D. contributed 30. Hartwig, J. W., Hu, H., Styborski, J. & Chung, J. Comparison of cryogenic flow to developing and defining the experimental test rig. boiling in liquid nitrogen and liquid hydrogen. Int. J. Heat Mass Trans. 88, 662–673 (2015). 31. Ganesan, V., Patel, R., Hartwig, J. W. & Mudawar, I. Review of databases and COMPETING INTERESTS correlations for saturated flow boiling heat transfer coefficient for cryogens in The authors declare no competing interests. uniformly heated tubes, and development of new consolidated database and universal correlations. Int. J. Heat Mass Trans. 179, 121656 (2021). 32. Mudawar, I. Flow boiling and flow condensation in reduced gravity. Adv. Heat Trans. 49, Ch. 5 (2017). ADDITIONAL INFORMATION 33. Chung, J. N., Dong, J., Wang, H., Darr, S. R. & Hartwig, J. W. Enhancement of Supplementary information The online version contains supplementary material convective quenching heat transfer of coated tubes by intermittent cryogenic available at https://doi.org/10.1038/s41526-022-00220-9. pulse flows. Int. J. Heat Mass Trans. 141, 256–264 (2019). 34. Chung, J. N., Darr, S. R., Dong, J., Wang, H. & Hartwig, J. W. Heat transfer Correspondence and requests for materials should be addressed to Jason Hartwig. enhancement in cryogenic quenching process. Int. J. Thermal Sci. 147, 106117 (2020). Reprints and permission information is available at http://www.nature.com/ 35. Xu, W., Cheng, C. & Zhang, P. Cryogenic flow quenching of horizontal stainless reprints steel tubes. Int. J. Heat Mass Trans. (2021). 36. Chung, J. N., Dong, J., Wang, H., Darr, S. R. & Hartwig, J. W. An advance in transfer Publisher’s note Springer Nature remains neutral with regard to jurisdictional claims line chilldown heat transfer of cryogenic propellants in microgravity using in published maps and institutional affiliations. microfilm coatings for enabling deep space explorations. Nat. Microgravity 7,21 (2021). 37. Hartwig, J. W., McQuillen, J. B. & Rame, E. Pulse Chilldown Tests of a Pressure Fed Liquid Hydrogen Transfer Line AIAA-2016-2186, AIAA SciTech Conference, January 4–8 (San Diego, CA, 2016). Open Access This article is licensed under a Creative Commons 38. Shaeffer, R., Hu, H. & Chung, J. N. An experimental study on liquid nitrogen pipe Attribution 4.0 International License, which permits use, sharing, chilldown and heat transfer with pulse flows. Int. J. Heat Mass Trans. 67, 955–966 adaptation, distribution and reproduction in any medium or format, as long as you give (2013). appropriate credit to the original author(s) and the source, provide a link to the Creative 39. Burgraff, O. R. An Exact Solution of the Inverse Problem in Heat Conduction Commons license, and indicate if changes were made. The images or other third party Theory and Applications. ASME J. of Heat Trans 86, 373–380 (1964). material in this article are included in the article’s Creative Commons license, unless 40. Kirillov, P. L. et al. The Look-Up Table for Heat Transfer Coefficient in Post-Dryout indicated otherwise in a credit line to the material. If material is not included in the Region for Water Flowing in Tubes FEI-2525. (Institute of Physics and Power article’s Creative Commons license and your intended use is not permitted by statutory Engineering, Obninsk, Russia, 1996). regulation or exceeds the permitted use, you will need to obtain permission directly 41. Leung, L. K. H., Hammouda, N. & Groeneveld, D. C. A look-up table for film boiling from the copyright holder. To view a copy of this license, visit http:// th heat transfer coefficients in tubes with vertical upward flow. In Proc. 8 Inter- creativecommons.org/licenses/by/4.0/. national Topical Meeting on Nuclear Reactor Thermal-Hydraulics, 671–678 (Kyoto, Japan, 1997). 42. Marquardt, E. D., Le, J. P. & Radebaugh, R. Cryogenic Material Properties Database This is a U.S. Government work and not under copyright protection in the US; foreign 11th International Cryocooler Conference, June 20–22, (Keystone, CO, 2000). copyright protection may apply 2022 npj Microgravity (2022) 33 Published in cooperation with the Biodesign Institute at Arizona State University, with the support of NASA

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