Hybrid cryogenic cooler for space flight applications
R. V. Annable
The hybrid cryogenic cooler is an intermittent Joule-Thomson refrigerator with a precooler in the form of apassive radiator. The properties of the J-T expansion and the gas storage vessel are used to select fluids onthe basis of available refrigeration per unit mass. Surface forces and container geometry are used to confineand control the liquid cryogen in a zero-gravity environment. The precooler and vaporized liquid are usedto reduce parasitic thermal inputs to the point where most of the heat of vaporization is available for usefulpurposes. Modifications can be made to increase the efficiency or extend the temperature range. Ambientstorage combined with efficient operation make the hybrid cooler attractive for space flight applications.
1. Basic Design and Operation
The hybrid cooler is a cryogenic refrigerator for sat-ellite applications that utilizes the Joule-Thomson ex-pansion of refrigerated gas. The basic design is shownin Fig. 1. High pressure gas from an ambient temper-ature (200C) storage vessel is pressure regulated andtransferred through a counterflow heat exchanger to aradiant cooler refrigeration stage.' The gas is cooledon this stage to the temperature range from 150 K to 170K (the exact temperature depending on the specificdesign). The refrigerated gas then passes through asecond counterflow heat exchanger and is expandedthrough a throttling (Joule-Thomson) valve into a lowpressure region on the second stage of the cooler. Theexpansion converts some of the gas into a liquid, theliquid temperature depending on the final pressure.The remaining gas then passes through the heat ex-changers and refrigeration stage and is exhausted atambient temperature. The exhaust is controlled by anabsolute pressure relief valve that thereby regulates thetemperature of the liquid in the second stage.
The refrigeration of the gas to the 150-170-K rangeprior to expansion greatly increases the fraction that isconverted to liquid. The basic system is then aHampson liquefier with a precooler, as described byScott.2 The second stage is surrounded by the radiantlycooled first stage (except for any optical ports to cooleddetectors) to reduce the parasitic thermal loads. Theconversion to liquid is continued until a given mass (say
The author is with ITT Aerospace/Optical Division, Fort Wayne,Indiana 46803.
a gram) has been stored in the second stage. The heatof vaporization is then available for refrigeration. Inaddition, the heat capacity of the vaporized liquid isavailable to offset, or even eliminate, the conductiveinput from the first stage produced by the mechanicalsupports and the second heat exchanger.3 Note thatduring this time there is no incoming gas to be cooledby the exhaust gas; the throttling process has beenturned off on the first stage. In addition, the radiativecoupling between stages can be made very small com-pared with other thermal loads by means of metallicradiation shields4 or conventional multilayer insulation.Under these conditions, nearly all the heat of vapor-ization is available for detector associated (optical andelectrical) thermal loads on the second stage.
When the liquid cryogen has been vaporized down toa predetermined level, the refrigeration system is turnedon again by the valve on the first stage. This valve islocated at the entrance to the heat exchanger that leadsto the Joule-Thomson valve. When the stored cryogenhas reached a second predetermined level, the valve isturned off. The gas capacity of the first stage must belarge enough to provide the refrigerated gas necessaryto produce the liquid cryogen contained between thetwo predetermined levels. In addition, the capacitymust be sufficient for the initial cool down of the secondstage. The duty cycle of the J-T expander is very low;a typical value is 0.1%.
Above a certain pressure (e.g., 18.6 X 106 N m- 2 fornitrogen), the fraction of gas converted to liquid (con-version factor) does not change significantly. The useof a pressure regulator set at this pressure then allowsthe components below the regulator to be designed foroperation at the control pressure rather than at thehigher maximum storage vessel pressure (e.g., 41.4 X 106N m-2). When the storage pressure has reached thecontrol pressure, a valve is turned on to bypass the
SU RFACE TENSION ILIQUID CRYOGENRETAINER ANDCAPACI TIV E MASSSEN SOR
Fig. 1. Schematic of the hybrid cryogenic cooler.
Fig. 2. Liquid conversion factor as a function of pressure at a re-frigeration temperature of 160 K. The expansions are to the normal
boiling points for 02 and N2 and to the triple point for CH4.
regulator and expand the remaining gas (down to theexhaust pressure) at a continuously decreasing con-version factor.
II. Conversion Process
The number of molecules per unit volume stored inthe high pressure vessel of volume V0 at a temperatureT. (assumed constant) and initial pressure P0 is givenby N/V. = P. (zkT.) -, where z is the compressibilityfactor5 and k is Boltzmann's constant. The corre-sponding gas density is given by pO = NM(VoNa)-l =PoM(zkToNa)-1, where M is the molecular weight andNa is Avogadro's number. The stored fluid producesa total refrigeration available at the second stage tem-perature that is equal to
Q = p0VoQve,
where Q = heat of vaporization at exhaust pressurePe,
= mean liquid conversion factor (g ofliquid/g of gas into second heatexchanger), and
e = conversion efficiency of second heatexchanger.
There is additional refrigeration available from theheat capacity of the vaporized liquid. This refrigerationcan be used to offset or eliminate the conductive inputthrough the second heat exchanger and its associatedsupport; it is given by
Q1 = p.V 0veupAT,
where cp = mean heat capacity at Pe and AT = tem-perature drop between cooler stages.
The liquid conversion factor x between a given inletpressure to the secoid heat exchanger and a given ex-haust pressure is given by6
x = (h4 - hl)/(h 4 - hliq),
where h = specific enthalpy,1 = high-pressure gas at inlet to exchanger,4 = low-pressure gas at exit from exchanger,
andliq. = liquid.
This equation follows from the fact that in a throttlingprocess the initial and final enthalpies are equal.7 Theconversion factors for oxygen, nitrogen, and methaneare shown in Fig. 2 as functions of the inlet pressure.8The behavior of ethylene is similar to that of methane;its conversion factor changes little with pressure as longas it is above the boiling pressure at the first (refriger-ation) stage temperature.
We have assumed that the temperature To of the
CU. N. TUBE FINS (HEAT EXCHANGER)
\ 9 ~~~~CRIESSOLDER ; , E X TPR
COLD TEMPERATURE SENSOR EPOXY GLASS
(SUPPORT B CONTAINERFOR HEAT EXCHANGER)
'- LIQUID SENSOR OUTPUT
(SECOND- STAGECYLINDER a COAXIALFIRST-STAGE CYLINDER)
storage vessel is constant. Only a very small amountof gas is used during a cycle of operation to reestablishthe level of stored cryogen. As a result, the corre-sponding cooling of the storage gas by VAP work is alsovery small. Moreover, the cycle is very long, so that thestorage gas will reestablish thermal equilibrium with thestorage vessel in the periods between gas usage.
The mass flow rate (g/sec) during liquid productionas well as the conversion factor (x) of gas into liquid arefunctions of the Joule-Thomson cryostat. A typicaldesign (from a laboratory model of an argon hybridcooler) is shown in Fig. 3. Note that the heat exchangeralso serves as the support for the second stage of thehybrid cooler. The heat exchanger consists of a heli-cally finned tubing coiled around a mandrel with sealingstrips (polyester thread) to prevent shell-side bypassing.It is commonly used in Joule-Thomson cryostats9 "10 andin some helium refrigerators." The design is derivedfrom the Giauque-Hampson and Parkinson heat ex-changers.12
For our purposes, the heat exchanger efficiency isgiven by e = (actualx)/(idealx), and in terms of ther-modynamics or heat transfer, the efficiency is given byE = (h2 - h1)/(h2 - hi), where the subscript 2 denotesthe high pressure fluid just before expansion, and theprime denotes the actual value. As an example, con-sider the use of argon at T, = 160 K and Pi = PC = 1.034X 107 N m-2 . We then have specific enthalpies (in Jg'-) of'3 h, = 172, h2 = 151.5, h3 = 233, hliq. = 74, andh4 = 275, so that x = 0.5124 for a perfect exchanger.(The subscript 3 denotes the unliquefied portion of theexpanded fluid.) If the conversion efficiency e is only0.95, the specific enthalpy of the exit gas is reduced toh = 265. The corresponding thermodynamic effi-ciency can be determined by means of the equation (h,- h'2) = (1 - x')(h4 - h3) that equates the heat given upby the incoming stream with that gained by the outgoingstream. We then have h = 156 and E = 0.80, that is,a liquid conversion efficiency of 0.95 requires only a verypoor heat transfer efficiency of 0.80. If we wish to in-crease e to 0.99, we find that an E value of 0.96 isneeded.
We can calculate the flow rate through the Joule-Thomson value from the empirical relationship ofReynolds or Perry.14 The relationships are for air andinclude functions of the ratio r of exhaust to inletpressure (Pe/Pc). We may expect the equation to hold,at least approximately, for gases similar to air, such asN2 , 02, and A, if we correct for the difference in molec-ular weight. This calls for multiplying by (M/29)1/2,where 29 is the molecular weight of air. Also, in ourapplication r is much less than one. If we now expresspressure in N/M2 , area in cm2, and temperature in K,we obtain w = 6.34 X 10-4PcAM12T-lI 2 g/sec, where Ais the area of the orifice. Using the equation for gasdensity given above (at PC and T2), the correspondingvolumetric flow becomes' 5 v = 0.852PAM-1/2T 2 stdliters/min, where we have assumed a compressibilityfactor of 1. For our argon experiment, we have A =7r(0.010) 2 /4 cm2 , M = 39.944, and T2 = 148 K, so that w= 0.267 g/sec and v = 9.00 std liters/min.
The dependence of w or v on PC, A, T2 , and M can beobtained readily from a simple physical model. A col-umn of gas of length that passes through an orifice ofarea A receives energy of force times distance, or PcAl,which is transformed into kinetic energy, 1/2 mV2 , on thelow-pressure side of the opening. The mass m equalspAl, where p is the density of the gas at a pressure P,and a temperature T2 . The linear velocity of the gas onthe low-pressure side is then V' = (2Pl/p) , and thevolumetric flow is v = V'A = A(2P0 /p)1/2 . Using theexpression for p and assuming a z value of one, this be-comes v = 1.289 X 104AT'/ 2 M-1/2 cm3 /sec or v = 2.085PcAM-1/2T2 I std liters/min. The ratio of the em-pirical equation to this ideal is 0.41. This factor is calledthe discharge coefficient and is the product of an arealcontraction and a velocity reduction in the actual fluidstream. 16
We are now in a position to determine the duty cycleof the J-T expansion, which is given by f = bb(Qvwxe)-1, where 'bb is the heat load on the second stageof the cooler. As an example, consider our laboratorymodel, in which 4 b = 0.025 W, Qv = 161.7 J g-1 (argonat 1 atm), and e - 0.99, so that f 1.14 X 10-3. Thisresult illustrates the low duty cycle nature of theJoule-Thomson expansion within the hybrid cooler.
Ill. Pressure Vessel Design and Fluid SelectionAs shown by the following analysis, the important
property of the pressure vessel material is the ratio ofits allowable stress s to its density p. This ratio is pro-portional to the ratio of yield strength a to density p,where the proportionality constant is the design margin.The alp ratio is called the specific strength. A com-parison of materials shows that among metals, titaniumalloys have the highest specific strength.1 7 However,filament-wound reinforced plastics have even higherspecific strengths, by a ratio of about 3 (4310/1400). Apractical compromise is available in the form of a tita-nium lined composite.
In general, we are interested in maximizing the re-frigeration we can obtain from a unit weight (or mass)of the filled storage vessel. The mass of gas in aspherical vessel of inside radius r is given by
m0 = %lrr3p, = V0 p.
For a thin-walled sphere, the mass of the container isgiven by
m = 47rr2 tp0 = 3V.(t/r)p,
where t is the wall thickness and PC the density of thematerial. But for a thin-walled sphere we also have'8tir = P/2s. The ratio of container mass to gas mass isthen m,/m 0 = 3Pop,/2sp0 .
If we use the equation for pO given in Sec. II, the ratiobecomes m,/m 0 = 3pczkToNa/2sM. For Pc in g/cm3 ,To2 in kelvins, s in N/M2, and M in g/mole, we obtainm,/m, = 3pzT 0/2.41 X 10-7sM = azTOjM. For manygases, the compressibility z is only a weak function ofpressure (up to at least the order of 4 X 107 N m-2 ). Inthis case, the mass ratio is very nearly independent ofboth the volume of the container and the pressure of the
* For To = 295 K; z = 1.07(02), 1.26(N2 ), 1.0(A), 1.04(CH 4 ),1.00(C2H 4); z = 1.05 assumed for F2, CO, and CF4 : z values at Po =4 X 107 N m- 2 (400 atm).
Table I. Mean Liquid Conversion Factors"
Fluid p b Pc b c
02 4.14 X 107 1.86 X 107 0.453N2 4.14 X 107 1.86 X 107 0.346A Constant 1.03 X 107 0.512CH4 4.14 X 107 1.03 X 107 0.672C2H4 4.14 X 107 d 0.854
a Refrigeration temperature of 160 K, except C2H 4 (158 K).b Pressures in N m-2.c To liquid at normal boiling point, except C2H 4, which is to liquid
at 123 K.d Above the boiling pressure at the refrigeration temperature (5.0
X 104 N m- 2 at 158 K).
gas. That is, the mass of the container, as well as themass of the gas, is proportional to the volume times thepressure. This means that the weight of the storagevessel for a given amount of gas is the same no matterhow the gas is stored, e.g., in one large container orseveral small ones, one small container at a high pres-sure or one large container at a low pressure. An im-portant exception is ethylene; its compressibilitychanges significantly with pressure, and there is amass-volume tradeoff.
In the case of a titanium sphere of the design de-scribed by Hughes and Herr 9 we have19 s = 4.05 X 108N m- 2 and c = 4.43 g/cm3 , so that a = 0.136. If afilament-wound reinforced plastic is used in place of thetitanium, we obtain a = 0.0442; and for the titanium-lined composite, a = 0.0676.
The ratio of refrigeration Q0 available at the secondstage to the mass m of the storage vessel plus the storedgas is given by
Qo/m = QYe(l + mc/mo)-l = Qve(1 + azTo/M)'1,
where we have used the equation for QO given in Sec.II. The ratio Q0 /mxe is shown in Table I for variousfluids and for titanium and filament-wound storagevessels. The upper and lower temperatures are theapproximate triple points and normal boiling points,respectively. The fluids were selected from the listgiven by Scott.2 0 We eliminated all fluids that haveinversion2 1 or critical 2 2 temperatures above normalambient or triple points above 105 K. Nitrogen fluoride(NF3 ) was also eliminated because of insufficient in-
formation on its thermodynamic properties and itsnonavailability from gas suppliers. Silane (SiH4) waseliminated because it is spontaneously flammable oncontact with air.
To complete our evaluation, we need to determine themean liquid conversion factor x for all the fluids on ourlist. We have done this for five of the fluids, as shownin Table II.8 On the basis of our (limited) information,we have selected the following fluids for operation be-tween 55 K and 125 K:
55 K to 90 K: oxygen (02);91 K to 103 K: methane (CH 4);104 K to 125 K: ethylene (C2H 4).
Argon (A) is nearly as good as oxygen and because it isinert may be preferred within its limited temperaturerange. Tetrafluoromethane (CF4) is inferior to meth-ane and ethylene even at a mean conversion factor ofunity. And finally, the toxicity of fluorine (F2) andcarbon monoxide (CO) tends to eliminate them fromconsideration.
IV. Zero-Gravity Operation
In order to operate in a zero-gravity environment,we must solve the problems of contact (of cryogen andcontainer), confinement (of the liquid to its storagevessel), and control (of temperature and amount).Appropriate forces must be employed to control theliquid and thereby solve the problems. The availableforces are surface (capillary), mechanical, electrostatic,and thermal. To start, we have excluded the use ofthermal forces; these would consume some of the re-frigeration and tend to disrupt the thermal contactbetween the cryogen and its container. Mechanicalforces are limited to those originating in either thespacecraft (from its spinning or attitude control system)or the expansion of the gas from the Joule-Thomsonnozzle. In general, our approach is to use surface forcesto control the configuration of the liquid23 so that it islocated during the long quiescent off period in a positionwhere thermal contact is achieved, and both its tem-perature and quantity are readily measured and con-trolled.
In utilizing the surface forces, we are in effect makinguse of the natural configuration of the liquid in itscontainer under conditions of reduced gravity. More-over, this configuration can be controlled by means ofthe geometry of the container, i.e., its bounding surfaceplus any internal surfaces. For example, a simple ge-ometry can be used in which the natural position of thecryogen under low-gravity conditions is between theplates of a measuring capacitor. In this way, we canemploy a capacitive technique for sensing the quantityof cryogen and controlling the operation of the Joule-Thomson on/off valve. However, the configuration ofthe liquid is disturbed and/or its initial placement de-termined when the Joule-Thomson control valve isturned on for the production of liquid. Our objectivehere is to expand the gas and inject the liquid into thecontainer in such a way that the cryogen quickly as-sumes its static (off) configuration when the valve is
turned off. This would seem to require that the newliquid not contact the old until it has dissipated all ormost of the kinetic energy from the expansion.
The condition of the cryogen during its vaporization(when the Joule-Thomson gas system is off) is quasi-static because of the very low rate of heat transfer andtherefore of liquid depletion. The configuration of theliquid under these conditions can be determined bymeans of drop tower tests or by the simulator describedby Olsen.24 In fact, our geometry may take the formshown by Olsen as an internal baffle (simulated cylin-der) in a rectangular (simulated cylindrical) container.The resultant liquid configuration is shown in Fig. 4,which illustrates a flight geometry for the Joule-Thomson stage. Sufficient liquid is produced duringthe on period of liquid production to fill the volumebetween the cylinders (the electrodes of the measuringcapacitor). The minimum amount, i.e., the level atwhich the Joule-Thomson system is turned on, may beset by the requirements of either the capacitance mea-surement or the configuration control. Based on theresults given by Olsen, it appears that the minimumquantity of liquid should fill at least 10% of the con-tainer volume. At this level, the liquid assumes a po-sition in the corners, in the manner indicated in Fig. 4.At this point, the capacitance measurement tends to beinsensitive to the presence of the liquid.
The position of the liquid cryogen within its containeris determined by the fact that it assumes a configurationof minimum energy. Under static (equilibrium) con-ditions, the energy consists of surface energy and gra-vitational (potential) energy. (Viscous energy can beneglected.) As the ratio of gravitational to surface en-ergy decreases, the liquid assumes a configuration ofminimum surface energy. This generally means thatthe surface area between the liquid and its vapor is aminimum. Thus, the liquid in the container of Fig. 4crawls into the space between the cylinders in order tominimize its exposure to the vapor. Operation of theJoule-Thomson gas system, of course, disturbs thequasi-static (equilibrium) condition that prevails duringthe off period of liquid vaporization. In particular, ifthe liquid from the Joule-Thomson nozzle is directedat the cryogen remaining in the container, a largeamount of kinetic energy is added to the surface energy.As a result, the previous configuration will no longer bethe stable one, i.e., the condition of minimum energy.For this reason, we propose that the store of cryogen leftin the container at the time of Joule-Thomson turn onbe kept in an undisturbed state and that the new liquidproduced by expansion not be combined with this storeuntil it has largely dissipated its kinetic energy.
To this end, we can remove as much energy as possi-ble from the incoming gas by precooling it to liquidtemperature prior to passage through the orifice. Thisapproach also eliminates the possibility of new liquidbeing entrapped in the gas from the expansion andcarried out the exhaust port. In addition, we can directthe Joule-Thomson nozzle to give a tangential velocityto the cryogenic fluid, thereby forcing the liquid to re-main in contact with a wall of the container until its
kinetic energy is dissipated. When the new liquidceases its spinning, it should be in a position to combinewith the remaining store of cryogen and assume thedesired stable configuration of the off period. Onemeans of doing this is shown in Fig. 4. Here, the newliquid spins against the outer cylindrical wall above themeasuring capacitor so that it enters the static con-finement region on the side nearer the Joule-Thomsonheat exchanger. Another means is to spin the newliquid around the inside of the inner cylinder, so that itenters the stable confinement region from the oppositeend. Because the liquid is initially outside the confinesof the measuring capacitor, the latter means would tendto overfill the storage volume.
Finally, a phase discriminator between the heat ex-changer and the liquid storage region can be added toensure that only vapor is vented out the exhaust port.In Fig. 4, the discriminator is a surface tension (passive)device in the form of a hemispherical dome over thecylindrical storage region. This dome tends to fill withvapor at all times and thereby blocks the passage of anyliquid that may try to leave the confines of the coaxialstorage volume.
V. Design Features
The design features of the hybrid cryogenic coolermore fully explain its operation and point out the de-partures from a conventional Joule-Thomson cryo-stat.
A. Operation Below the Normal Boiling Point
As the exhaust pressure is reduced below 1 atm, thetemperature is reduced below the normal boiling point.At the same time, the heat of vaporization Q, is in-
H H
VAPOR o N0
A A
C1 C2 CZA 1 C C2 C2A 'C
CI~~~~ ~~~~~~~~~~~~~~~ CI )ICA1 A '.' A A
.1 0 0 0o
(a) ~~~~~~~(b)Fig. 4. Liquid confinement in a zero-gravity environment: (a)During cryogen boil-off (Joule-Thomson off); (b) during cryogenproduction (Joule-Thomson on). A = liquid-vapor interface atmaximum liquid mass; B = liquid-vapor interface at minimum liquidmass; 0 = opening for passage of liquid; C1, C2 = outer and innercylinder, respectively, of surface tension retainer/measuring capacitor;I = electrical insulator; D = phase discriminator (hemisphericaldome); H = Joule-Thomson heat exchanger; N = Joule-Thomson
nozzle; L = liquid stream; K = cooling coil for incoming gas.
creased, and the mean conversion of factor is de-creased such that QY, the refrigeration produced, islittle changed. For example, the calculations for anoxygen system show that going from the normal boilingpoint down to the triple point decreases the refrigera-tion by only 7%. Moreover, nearly all the refrigerationis available for the detector related thermal load, i.e.,it is nonparasitic. As a result, one may operate attemperatures below the normal boiling point at onlysmall penalties in terms of refrigeration and heatload.
B. Dual Use of the Heat ExchangerThe heat exchanger between the radiant cooler (re-
frigeration) stage and the Joule-Thomson stage servestwo distinct functions:
(1) As a conventional J-T heat exchanger duringliquid formation to transfer heat from the incoming gasstream to the outgoing gas stream.
(2) As a heat exchanger during boil-off to transferconductive heat coming from the first stage to the ventgas going back to the first stage.
The design used is that of a conventional J-T system(Sec. II), although modification may be necessary whena low exhaust pressure is required (e.g., for 02 or C2H4at the triple point).
C. Dual Benefits from the First Stage of CoolingThe first or refrigeration stage of cooling serves two
purposes:(1) It increases the fraction of gas converted-to liq-
uid.(2) In combination with vent gas cooling from the
cryogen boil-off, it reduces parasitic heat loads (con-ductive and radiative) on the second stage to a very lowvalue.
Table III shows the increase in x when the gas iscooled from 295 K to 160 K. Results are given for 02,A, and N2 at an exhaust pressure of 1 atm.
D. Joule-Thomson ValveAccording to Buller,'0 the choice of the Joule-
Thomson orifice size and the operating pressure in aconventional J-T system involves a trade-off amongcool-down time, operating time, and efficiency of heatexchange. Thus, a short cool-down calls for a largeorifice and a high pressure. This increases the effi-ciency of heat exchange but limits the operating time.One result has been the development of self-regulatingor throttling cryostats, as described by Buller. If thecool-down time is not critical, e.g., as in the MarinerMars 1969 ir spectrometer, 9 the orifice can be made verysmall while the high pressure is maintained. The limitin this direction is set by plugging of the orifice bycontaminants (particles and condensables).
In the case of the hybrid cooler, however, we are notconstrained by requirements on the cool-down time.We can use a relatively large orifice and the higherpressures desirable for good heat exchanger efficiencyand a high liquid conversion factor. The Joule-Thomson portion of the hybrid cooler operates in an
Table . Increase In Conversion Factor With Cooling
02 at10.3X AatlO.3X N2 atl2.1XT2,4 106 N m- 2 106 N m- 2 106 N m- 2
295 K 0.551 0.0664 0.0594160 K 0.488 0.512 0.358
intermittent or cyclic fashion and not in the continuousfashion of a conventional cryostat. The high flow ratesduring liquid production not only increase the efficiencyof the heat exchanger, they also keep the duty cycle (ontime of the control valve) low.
We may therefore conclude that the hybrid cooleravoids the constraints and problems associated withboth the cool-down time and the plugging of a Joule-Thomson system. The orifice and pressure (andtherefore flow rate) can be selected on the basis of heatexchanger efficiency and the liquid conversion factor(i.e., on the basis of useful refrigeration produced). Thepotential for plugging is further reduced by the refrig-eration of the gas on the radiant cooler stage. This, ofcourse, removes condensable material prior to enteringthe Joule-Thomson stage.
VI. Modifications
There are several ways in which the hybrid cooler canbe modified to increase its conversion efficiency or toreduce its operating temperature. These include theuse of multiple stages and the use of the cryogen in itssolid form.
A. Two Stages of Refrigeration
The conversion factor is increased by a second stageof refrigeration, that is, by adding a Joule-Thomsonstage to a two stage radiant cooler. Such a design wouldprobably only be appropriate for a relatively large ra-diant cooler. To get some idea of the increase in x, wewill compare two oxygen systems, one operating froma single stage radiant cooler at 163 K and one operatingfrom a double stage radiant cooler at 123 K. We haveassumed a control pressure of 150 atm (15.2 X 106 Nm-2) and an exhaust pressure of 1 atm (so that theoxygen is at its normal boiling point of 90.2 K). Theresults are25 x = 0.468 at T1 ,4 = 163 K and x = 0.743 atT1,4 = 123 K. The production of liquid is increased by59% when the gas is cooled to 123 K instead of 163 K.Or, viewed another way, the weight of the filled gasstorage vessel can be decreased by a factor of 0.63X.
B. Two Joule-Thomson StagesLower temperatures can be reached by adding a
second Joule-Thomson stage. An example is a nitro-gen/hydrogen system; a conventional system of thiskind is described by Hughes and Herr.10 If there areno detectors on the first J-T stage, its heat load will beextremely low. Under this condition, the stored gasrequired for the first expansion is very small, and thesize and weight of the storage vessel are determined bythe gas needed for the second expansion.
a Gas -stored at minimum compressibility pressure 2 7 (z = 0.34 atP = 1.0 X 07 N m 2 and T = 295 K).
The fluids with boiling points below the triple pointof oxygen are neon, hydrogen, and helium. Appropriatecombinations are neon/oxygen, neon/nitrogen, andhydrogen/nitrogen. A system that goes down to thenormal boiling point of helium (4.2 K) would be veryinefficient because of its low heat of vaporization (20.6J/g) and its very low conversion factor for an expansionfrom the 55-65-K range. The use of hydrogen and neonwould allow us to reach temperatures in the 14-27-Krange. Even then, the conversion factors are relativelylow (e.g., 0.253 for hydrogen at its normal boiling point,starting with liquid nitrogen at 65 K), and such systemswould generally be limited to relatively low heatloads.
C. Other Forms
It is possible, of course, to combine the above twoforms for both greater efficiency and a reduction intemperature. In addition, the hybrid cooler can comein at least two other basic modifications: a two-phasecooler and a two-fluid cooler.
In the first form, the liquid is produced at a highertemperature and pressure and then converted into solidat a lower temperature and pressure after the J-T gassystem has been turned off. The zero-gravity con-finement problem is largely eliminated in this form; themain consideration is then to control phase (liquid/gas)separation during operation of the J-T valve. Thelower temperature limit is set by the pressure dropacross the exhaust side of the heat exchanger, which canbe made 1.0 Torr (130 N m- 2 ) or less with no significantreduction in the conversion efficiency. At this pressure,we can use solid methane down to approximately 65 K,solid argon to 37.5 K, solid hydrogen to 9.5 K, and liquidhelium to 1.3 K.
In the second form, two fluids are used in the sameJoule-Thomson expansion. Such a form could be usedto establish a triple point temperature below that ofeither component or to create a solid/liquid mixture(slush). For example, a binary system of nitrogen andoxygen has a triple point at about 50.2 K at a level of221/2% nitrogen.26 Slush mixtures (with various ratiosof liquid to solid) can be formed between this temper-ature and the triple point of oxygen (54.35 K) or thetriple point of nitrogen (63.15 K).
VIl. Operational Characteristics
The hybrid cryogenic cooler has several characteris-tics that make it attractive for satellite-borne applica-tions:
(1) The cooler has a large refrigeration efficiency interms of the ratio of usable (available) thermal load tototal thermal load.
(2) There is no stored cryogen when the cooler is notoperating. The system can be launched at ambient,heated for outgassing, and reheated (if necessary) fordecontamination.
(3) Power consumption is very small. It is limited tothe level necessary for sensing and control; no electricalpower is dissipated as part of the refrigeration pro-cess.
(4) The consumption of the refrigeration capacity canbe delayed or interrupted.
(5) Because the radiantly cooled first stage operatesat a relatively high temperature (150-170 K), the designis not sensitive to the particular spacecraft and itsorbit.
The refrigeration efficiency (1) results in a lowerweight cooler and reduced size. The second feature (2)eliminates the ground handling and launch logisticsproblems associated with stored cryogen coolers.Feature (4) means the hybrid cooler is useful for mis-sions with long delays or with intermittent usage (i.e.,with long nonoperating periods) with little or no re-duction in useful operating time. Because of the lastfeature (5), the hybrid system can be used on missionswhich do not provide the cold space view necessary forcryogenic (<125-K) passive radiant coolers.
Another useful feature of the hybrid cooler is thechoice of working gas to provide the maximum refrig-eration at a given temperature. This is shown in TableIV, which is based on Tables II and III. The reciprocalrelationship between temperature and refrigeration isquite apparent. Also shown in Table IV is the increasein the refrigeration to mass ratio of ethylene obtainedby reducing the storage pressure (below 4 X 107 N m- 2 ),a change that decreases the compressibility factor. Thepenalty for this improvement is, of course, an increasein the volume of the storage vessel. Mass reduction canalso be realized at lower temperature by means of atwo-phase cooler. For example, the replacement ofliquid oxygen at 65 K by solid methane at the sametemperature reduces the mass by a factor greater than2 when the first stage is at 160 K.
The development of the hybrid cryogenic cooler asdescribed in this paper was supported by the ITT in-dependent research and development program. Thework was performed in the Electro-Optical SystemsDepartment of the Aerospace/Optical Division of ITTunder the management of D. J. Juarez.
1. R. V. Annable, Appl. Opt. 9,185 (1970); R. D. Hudson, Jr., and J.W. Hudson, Eds., Infrared Detectors (Wiley, Halstead Div., NewYork, 1975).
2. R. B. Scott, Cryogenic Engineering (Van Nostrand, Princeton,N.J., 1959), Sec. 2.3.
3. A. Bejan and J. L. Smith, Jr., in Advances in Cryogenic Engi-neering, Vol. 21, K. D. Timmerhaus and D. H. Weitzel, Eds.(Plenum, New York, 1976), paper G-1.
4. R. V. Annable, Appl. Opt. 15, 1860 (1976).5. R. W. Vance and C. H. Reynales, in Applied Cryogenic Engi-
neering, R. W. Vance and W. M. Duke, Eds. (Wiley, New York,1962), Chap. 2.
6. Ref. 2, Sec. 2.3. Note also that (h4 - hijq) = Qv + UpAT.7. R. W. Zemansky, Heat and Thermodynamics (McGraw-Hill,
New York, 1951), Sec. 11.10.8. The conversion factors were calculated from the data contained
in entropy-temperature and enthalpy-pressure charts suppliedby the NBS Cryogenics Div., Institute for Basic Studies (Boulder,Colo. 80302).
9. J. L. Hughes and K. C. Herr, Cryogenics 513 (Sept. 1973).10. J. S. Buller, in Advances in Cryogenic Engineering, Vol. 16, K.
D. Timmerhaus, Ed. (Plenum, New York, 1973), paper E-7.11. R. N. Meir and R. B. Currie in Advances in Cryogenic Engi-
neering, Vol. 13, K. D. Timmerhaus, Ed. (Plenum, New York,1968), paper G-3.
12. Ref. 2, Sec. 2.11.13. Temperature-Entropy Chart for Argon (NBS Cryogenics Div.,
Institute for Basic Studies, Boulder, Colo. 80302), Chart D-61.
14. L. S. Marks, Ed., Mechanical Engineer's Handbook (McGraw-Hill, New York, 1951), p. 2080.
15. Application of the equation to the nitrogen circuit of the Hughesand Herr experiment (Ref. 9; T1 = T4 = 295 K, P = 1.86 X 107N m- 2, A 2.03X 10-5 cm2, M = 28, and T2 165 K) yields v =
4.73 std liters/min, in good agreement with the observed flow of4.5-5.0.
16. Ref. 14, p. 239.17. Materials Selector 75 (Reinhold, New York, Sept. 1974), p. 6.18. R. J. Roark, Formulas for Stress and Strain (McGraw-Hill, New
York, 1965), p. 299.19. Metals Handbook, Vol. 1, Properties and Selection (American
Society for Metals, Metals Park, Ohio, 1961).20. Ref. 2, p. 268.21. He, H2 , and Ne.22. C2 , H6, and CClF3.23. M. W. Dowdy et al., Surface Tension Propellant Control for
Viking 75 Orbiter, AIAA Paper 76-596, presented at AAIA/SAE12th Propulsion Conference, Palo Alto, Calif., 26-29 July 1976.
24. W. A. Olsen, Simulator for Static Liquid Configuration in Pro-pellant Tanks Subject to Reduced Gravity, NASA TND 3249,Lewis Research Center, May 1966. The simulator is also de-scribed by C. L. Strong, in The American Scientist column in Sci.Am. 106, 226 (1972).
25. Based on enthalpies in Chart D-10, NBS Cryogenics Div., Insti-tute for Basic Studies (Boulder, Colo. 80302), reprinted from W.H. Keesom et al., Communs. Kamerlingh Onnes Lab., Suppl. No.112d (U. Leiden, 1955).
26. Ref. 2, p. 286.27. W. E. Forsythe, Smithsonian Tables (Smithsonian Institution,
