liq

Ch, United, Un

Keywords:

arisentersre c

contact along the pipe walls. The lower kinematic viscosity and surface tension of LH2, along with

ve beeod indbe use

exible, robust systems can be conceived, fundamental under-standing of the uid mechanics and heat transfer is rst required.

To enable all of the aforementioned systems, efcient methodswith which to transfer cryogenic propellant are required in orderto minimize consumption and loss. For example, before restartof an in-space cryogenic engine [28,29,35], or before propellant

liquid may ow,fore necessboiling an

The purpose of this paper is to compare the fundaphysics of two different sets of chilldown experiments conat two different Reynolds (Re) regimes using two differento understand how chilldown is affected by ow regime andthermophysical properties. The paper is structured as follows: abrief review of relevant literature is presented, along with back-ground into cryogenic transfer line chilldown. Next, descriptionsof both low and high Re number experimental hardware are given.Then experimental results are compared between systems.Temperature traces and ow visualization, along with computed

Corresponding author.E-mail address: Jason.W.Hartwig@nasa.gov (J. Hartwig).

International Journal of Heat and Mass Transfer 88 (2015) 662673

Contents lists availab

International Journal of H

.ecostly insulation systems and boil-off of precious propellant. Thelow normal boiling point (NBP) and low surface tension make itdifcult to transfer single phase liquid from a storage tank, makingvapor ingestion downstream highly probable. Therefore, before

Before continuous, steady, vapor-free propellantthe transfer line must rst be chilled down, therethe desire to understand the underlying owtransfer associated with chilldown.http://dx.doi.org/10.1016/j.ijheatmasstransfer.2015.04.1020017-9310/Published by Elsevier Ltd.itatingd heat

mentalductedt uidsmance in-space engines, in-space fuel storage depots, in-spaceresource utilization systems, cooling, refrigeration, liquefaction,thermal management, fuel cells, and life support systems.However, there are challenging aspects when working with cryo-genic liquids such as liquid hydrogen (LH2) due to the thermophys-ical properties. For example, the low normal boiling point makesLH2 particularly susceptible to parasitic heat leak, resulting in

pellant tank and transfer line hardware down to cryogenic temper-atures so that vapor-free liquid may ow between two points.Chilldown therefore represents the rst stage of cryogenic uidtransfer. In-space engines require vapor-free propellant ow toavoid combustion instability issues during restart, and in-spacedepots will require vapor-free liquid ow in order to achieve veryhigh liquid ll fractions in the customer spacecraft receiver tank.Nucleate boilingTransition boilingChilldownLiquid nitrogenLiquid hydrogenCryogenic heat transfer coefcient

1. Introduction

For decades, cryogenic uids hachemical, aerospace, medical, and fopellant technology development canreduced parasitic heat leak and higher Re numbers relative to the LN2 experiments, causes chilldownto proceed almost immediately into the nucleate boiling regime, in comparison to low Re ows where>75% of the chilldown is spent in vapor lm boiling.

Published by Elsevier Ltd.

n used throughout theustries. Cryogenic pro-d to enable high perfor-

transfer from the in-space storage depot tank to the customerreceiver tank [38,21], the transfer line connecting the propellantstorage tank and engine or the line connecting depot storage andcustomer receiver tanks must be rst be chilled down quicklyand efciently. Chilldown is dened as the process of cooling pro-Available online 20 May 2015traces and ow visualization indicates that the chilldown process evolves much more rapidly at higherRe numbers due to a quick transition from vapor ow to annular liquid ow and near immediate liquidComparison of cryogenic ow boiling inhydrogen chilldown experiments

Jason Hartwig a,, Hong Hu b, Jeremy Styborski c, J.N.aNASA Glenn Research Center, Propulsion and Propellants Branch, Cleveland, OH 44135bMechanical and Aerospace Engineering, University of Florida, Gainesville, FL 32611, Unc Turbine Durability, Hot Section Engineering, Pratt & Whitney, East Hartford, CT 06118

a r t i c l e i n f o

Article history:Received 7 December 2014Received in revised form 29 April 2015Accepted 29 April 2015

a b s t r a c t

This paper presents a compfer line chilldown experimconducted at low Re numbheat transfer coefcients a

journal homepage: wwwuid nitrogen and liquid

ung b

ited StatesStatesited States

on between experimental results from recent liquid hydrogen (LH2) trans-s at high Reynolds (Re) numbers versus liquid nitrogen (LN2) experiments. Parasitic heat leak, inner wall temperatures, inner wall heat uxes, andomputed to compare between the two systems. Analysis of temperature

le at ScienceDirect

eat and Mass Transfer

l sevier .com/locate / i jhmt

Heaparasitic heat leak, inner wall temperatures, inner wall heat uxes,and heat transfer coefcients are used to compare chilldown char-

Nomenclature

A area [m2]cP specic heat of transfer line [J/kg K]d pipe diameter [m]dtchill time interval to steady state chilldown [s]F view factor, dimensionlessG mass ux [kg/m2 s]g gravity [m/s2]h enthalpy [kJ/kg]hi heat transfer coefcient [W/m2 K]k thermal conductivity [W/m K]L characteristic length scale [m]m mass of transfer line [kg]_m mass ow rate [kg/s]P pressure [Pa]Qow ow energy [J]Qline stored line energy [J]Qparasitics parasitic heat leak energy [J]_Qgascond gas conduction heat leak [W]_Qrad radiation heat leak [W]_Qsensor sensor heat leak [W]_Qsolidcond solid conduction heat leak [W]q00axial axial heat ux [W/m

2]q00i inner wall convective heat ux [W/m

2]q00w heat ux at the wall [W/m

2]r radius [m]Re Reynolds number

J. Hartwig et al. / International Journal ofacteristics for the two different ow regimes. Finally, both data setsare compared with empirical correlations used to predict the crit-ical heat ux and Leidenfrost temperature.

2. Background and governing physics

The purpose of line chilldown is to cool hardware down to cryo-genic temperatures so that single phase cryogenic liquid may owbetween two points in a system. An energy balance for any transferline system, which includes all piping, valves, pumps, etc. can beexpressed in the following form:

Qflow Qline Qparasitics 1where Qline is the thermal energy stored in the line:

Qline Z tsst0

mcPdTdt

dt 2

Qparasitics is the sum of total unwanted heat leak into the system,which includes radiation, solid and gaseous conduction (assumingsystem is in a partial or full vacuum environment), and heat leakdue to instrumentation:

Qparasitics Z tSSt0

_Qrad _Qsolidcond _Qgascond _Qsensor

dt 3

where

_Qrad X

erAiFij T4i T4j

4

_Qsolidcond ksolidAc dTdz 5

_Qgascond kgas dTdx Asurface 6and _Qsensor is the measured electrical heat input from the pressuretransducers.

T temperature [K]t time [s]We Weber number, dimensionlessa thermal diffusivity [m2/s]c surface tension [N/m]e emissivity, dimensionlessg efciency parameter, dimensionlessl viscosity [Pa s]q density [kg/m3]r StefanBoltzmann constant [W/m2 K4]0 initial statec cross-sectionalCHF critical heat uxe exitfg uid to gasgas gasi inner surfacein inletj jth surfacel liquidLei Leidenfrosto outer surfaceSAT saturationsolid solid surfaceSS steady statev vapor

t and Mass Transfer 88 (2015) 662673 663Qow is the ow energy stored in the uid available to combatthe other two energy terms:

Qflow Z tSSt0

_mhe hin dt 7

For a subcooled liquid in the storage tank, both the latent andsensible energy is available for line chilldown:

Qflow Z tSSt0

_m hfg h hSATP

dt 8

By denition, in order to achieve perfect vapor-free liquid at the endof the transfer line, the subcooled portion (sensible energy in Eq.(8)) of the ow energy must exceed the sum of line and parasiticenergies absorbed by a uid element.

The combination of cryogenic uid ow and warm hardwareimplies that there will be two phase ow, vigorous boiling, andheat transfer as the chilldown process evolves in time. The uidmechanics and heat transfer of cryogenic transfer line chilldowncan be accurately mapped using a boiling curve, which is alog/log plot of heat transfer as a function of the temperatureabove saturation. Shown in Fig. 1 is a typical saturated boilingcurve for water. There are four distinct heat transfer regions sep-arated by three characteristic points. Because the system is ini-tially at a warm ambient temperature, which is signicantlywarmer than the cryogenic uid saturation temperature, chill-down generally begins in the vapor lm boiling regime, due toviolent ashing of the liquid. As the system chills, and as walltemperature decreases, the system approaches the Leidenfrostpoint, or point of minimum heat transfer between cold uidand warm pipe. Heat transfer is a minimum at the Leidenfrostpoint due to inefcient heat transfer between cold vapor andwall. The next region is known as the transition boiling regime,or regime between liquid dominated chilldown in nucleate

Hea664 J. Hartwig et al. / International Journal ofboiling, and gas dominated chilldown in vapor lm boiling. Asthe wall temperature decreases, the system quickly approachesthe critical heat ux, or point at which liquid begins to contactthe wall. The next region is characterized by liquid dominatednucleate boiling regime. As the wall further cools, and as thevoid fraction decreases, the system approaches the onset ofnucleate boiling (ONB) point, characterized as the point at whichthe system evolves from nucleate two phase cooling to singlephase liquid convection. Heat transfer is a maximum at thepoint of critical heat ux due to the highly efcient cooling pro-cess of boiling liquid through the use of sensible and latent heatand due to the fact that the insulating vapor layer is absentwhile the temperature difference between wall and uid is thegreatest. It is interesting to note that the evolution of chilldowngenerally mimics a reverse boiling curve. The Leidenfrost point,point of critical heat ux, and ONB are all characteristics pointsfor each uid, but the general shape of the curve is the same.

There are numerous studies on room temperature or storableuid ow boiling and heat transfer such as water [1820] and car-bon dioxide [23]. Previously reported relevant studies on cryogenicboiling, heat transfer, and chilldown are listed as follows: Westbyeet al. [39] conducted laminar and turbulent horizontal 1-g chill-down tests using R113. Kawanami et al. [16,17], Yuan et al.[41,42], Zhang and Fu [44], and Hu et al. [9,10], conducted varioushorizontal, upward, and downward 0 g and 1 g laminar and turbu-lent tests with LN2. Heat transfer coefcients for LN2 are availablein the literature for pool boiling [2,11], forced convection with

Fig. 1. Characteristic saturatedt and Mass Transfer 88 (2015) 662673subcooled liquid [36], forced convection with nitrogen at supercrit-ical pressures [37], and lm boiling [6]. Additional ow boilingcharacteristics are detailed in the recent papers, see for example[9]. Flow boiling information for liquid oxygen (LOX) is availablein [30].

Meanwhile, the only known relevant chilldown experimentusing LH2 was the work from [3], for horizontal 1-g turbulent ow.They did not however report information regarding ONB,Leidenfrost point, or ow boiling heat uxes. Pool boiling heatuxes are available in [32,7,24], as well as some ow boiling char-acteristics for LH2 are available in the literature [33,34]. Howeverthe heat ux data does not match the well-known theory from[27], which predicts that bulk uid velocity magnitude has littleeffect on the critical or maximum heat ux value; data in [34]show a dependence on velocity and do not converge.Additionally, two separate methods were used to deduce the massow rate, one using a weight scale, and one using the gas ow rateinto the liquid supply dewar. While the weight scale can be used asan accurate measure, provided that tests are run for a very longtime (long enough to establish a steady state velocity), using thegas ow rate into the liquid dewar introduces complications dueto condensation heat transfer across the tank liquid/vapor inter-face, due to compressibility effects, and due to potential solubilityof the pressurant gas into the liquid. Using the pressurant gas owrate is only accurate in the limit of an isentropic tank drain, whereheat transfer between pressurant gas and liquid cryogen isminimized.

chilldown curve for water.

3. Experimental setup

3.1. LN2 experiments

Low Re number liquid nitrogen line chilldown experimentswere performed at the University of Florida. Specic details regard-ing experimental hardware and procedure are available in [9].Fig. 2(a) and (b) shows a photo and schematic of the test setup,respectively. The system was designed to t inside an apparatusframe to minimize line lengths and thus parasitic heat leak asshown in Fig. 2(a). The test section was a 1 cm outer diameter,0.1 cm thick, 25.4 cm long see-through Pyrex section. The inletcondition was saturated ow, for quality between 0 and 0.122.Re number range was 10004000. Gaseous nitrogen was used topressurize and drain the LN2 storage dewar. A three way valve

was used to prechill a portion of the line leading up to the test sec-tion. The ow was then routed into a vacuum section which con-tained the sight glass used for ow visualization. Flow could berouted vertically upward or downward or horizontally. Flow wasthen routed through a series of heat exchangers designed to com-pletely vaporize the uid before entering the gas ow meter andvent line. A vacuum pump was used to pump down the pressureon the sight glass section to 23 kPa.

Temperature instrumentation is shown in Fig. 3. A set of threethermocouples were mounted at three different stations alongthe outer test section, spaced axially apart 120 to measure axialand longitudinal temperature variations as the system chilled in.Pressure was measured inside the liquid dewar and upstreamand downstream of the test section. A high speed camera was usedto visualize the ow to time correlate with the temperature and

J. Hartwig et al. / International Journal of Heat and Mass Transfer 88 (2015) 662673 665Fig. 2. (a) Photo and (b) test schematic for the low Reynolds number liquid nitrogen line chilldown test setup.

Fig. 3. Location of temperature sensors for the low Reynolds number liquid nitrogen experiments.

Hea666 J. Hartwig et al. / International Journal ofpressure data. Vacuum pressure was measured with a low pressurevacuum gauge. Mass ow rate was controlled by the pressurant gasow.

3.2. LH2 experiments

High Re number liquid hydrogen line chilldown experimentswere performed at the NASA Glenn Research Center SmallMulti-Purpose Research Facility (SMiRF). Specic details regardingexperimental hardware and procedure are available in [26] onlygeneral details required to understand the experimental resultsin the current work are presented here. A schematic of the testsetup is shown in Fig. 4. A 1.84 m3 (487 gallon) storage vesselwas used to store the LH2 and condition the liquid to the desiredinitial saturation temperature between 20.3 K < T < 24.2 K. A hori-zontal liquid acquisition device (LAD) located inside the tank was

Fig. 4. High Reynolds number liquid hydrogen line chilldown test setup.used to ensure liquid ow out of the bottom of the tank. Flowwas immediately routed through a Coreolis ow meter (FM),which was close coupled to the tank to minimize two phase owreadings. Liquid was then routed to a manifold, where a series ofvalves and orices were used to control ow rate. Flow was routedvertically upward through a dummy valve, which was used toadd mass to the transfer line. The ow control manifold could deli-ver LH2 ow rates over a wide range corresponding to a wide rangein liquid Re numbers (1.84 1044.33 105). The inlet conditionwas saturated liquid. The vertical upward direction was chosento mimic axisymmetric heat transfer into the pipe as in the micro-gravity pipe ow case [40,43]. Liquid then passed through a sightglass section for two phase ow visualization as the chilldown pro-cess evolved. Flow was nally routed up out of the top of the vac-uum chamber, through approximately 30 m of piping through aheat exchanger and eventually to the vent stack located outsidethe facility. The outer diameter of the transfer line was 1.27 cm(0.5 in) and the inner diameter was 1.02 cm (0.402 in).

The whole tank and line chill down assembly rested inside of avacuum chamber set to a background pressure of 1 106 torr and250 K to simulate a cold solar inertial orbit of Low Earth Orbit(LEO). The entire line assembly was constructed from 1.27 cm(0.5 in) outer diameter (OD) stainless steel (SS), while the test sec-tion for ow visualization was made of Pyrex glass. The portion ofthe line assembly between tank and valve manifold was always inliquid, since there was no control valve between LAD and lineassembly. This is deemed the cold portion of the line, while allpiping and components downstream of the control valves isdeemed the warm portion of the line. The length and mass ofthe cold portion of the assembly was approximately 2.34 m and5.1 kg while the warm portion including the two valves, piping,and sight glass was approximately 2.41 m and 2.3 kg.

Flow rate was measured using a Coreolis FM. Pressure was mea-sured in the tank, upstream of the manifold, and upstream anddownstream of the sight glass. Stream temperatures were mea-sured in the storage tank, upstream of the valve manifold (SD15),and downstream of the sight glass (SD23), while skin temperatureswere measured on the SS pipe (SD17, SD18) and Pyrex sight glasstube (SD19). A high speed camera was used to visualize ow tocorrelate with the pressure and temperature data. Uncertaintiest and Mass Transfer 88 (2015) 662673for all these measurements are listed in Hartwig et al. [8].

4. Results and comparison

4.1. Temperature traces and ow visualization

Fig. 5 plots a typical outer wall temperature versus time tracefrom the LN2 chilldown tests from [9]. Still images from the highspeed camera are shown to correlate with the temperature/timecurve. Also shown in the gure is the quench front movement dur-ing the transition boiling regime. As shown, at lower Re numbersand higher parasitics, most of chilldown is spent in the lm boilingregime. Then, a dispersed droplet ow comprised of small

HeaJ. Hartwig et al. / International Journal ofspherical liquid drops embedded in the vapor is observed. With afurther wall temperature decrease, the small droplets become den-ser and coalesce to form some larger and stretched droplets. Afterthat, a continuous liquid core occurs which ows in the center ofthe tube and is surrounded by a thin layer of vapor lm. This phe-nomenon is called the inverted annular ow. At this time, the lmis thick enough to separate the wall and liquid core efciently.With a further decrease of the wall temperature, the vapor lmbecomes thinner while the liquid core occupies more area. Whenthe wall temperature is reduced to the rewetting temperature,the liquid is able to come into direct contact with the tube wall.Some short liquid laments can be seen. The transient boiling,characterized by intermittent liquid-wall contact and violent bub-ble generation, is observed during a limited period. The quenchingfront occurs at the point of maximum heat ux. The vaporliquidinterface is highly wavy because of the KelvinHelmholtz instabil-ity. At this time the heat transfer rate from the hot wall to the liq-uid is signicantly increased and a rapid drop of the wall

Fig. 5. Outer wall temperature trace and ow visualization for liquid nt and Mass Transfer 88 (2015) 662673 667temperature slope can be observed. A rapid change from transitionow to the dispersed bubbly ow will occur directly after thequenching front. The prevailing boiling regime after the quenchingfront passes is the nucleate boiling with clear small bubbles gener-ated at the wall. Finally, the system achieves saturated LN2 temper-atures and bubbly ow.

Fig. 6 plots outer wall temperature versus time traces from theLH2 chilldown experiments for saturated 20.3 K liquid for a med-ium ow test (0.014 kg/s, Re = 1.22 105). Superimposed in theplot is the inner wall stream temperature reading immediatelydownstream of the sight glass as well as the liquid mass owmeterreading as a function of time. Also plotted are time correlated videoclips from the sight glass section. As shown, at higher Re numbersand lower overall parasitic heat leak, the outer wall skin tempera-tures immediately plummet within the rst 20 s of chilldown. Bothskin diodes bottom out in succession. Time correlated ow visual-ization reveals that this is due to liquid droplets forming along thewalls almost instantaneously from the start of chilldown. After the

itrogen line chilldown test for TSAT = 79 K, 0.0044 kg/s, Re = 3085.

Hea668 J. Hartwig et al. / International Journal ofskin diodes bottom out near the saturation temperature, themajority of the rest of the chilldown is spent transitioning fromannular to slug to bubbly ow. The stream temperature readingshown in red eventually reads saturated liquid hydrogen tempera-tures after 155 s. Slight, temporary increases in internal streamtemperature readings are attributed to transitions in ow prolesfrom vapor to droplet and from annular to churn/bubbly ow [26].

As shown in Fig. 6, most of the chilldown is spent in annular andchurn/bubbly ow as the system reaches an equilibrium betweenparasitic heat leak and ow energy. Since parasitics for this exper-iment were slightly higher than maximum ow energy, perfect

Fig. 6. Outer wall temperature trace, stream temperature, and ow visualization fort and Mass Transfer 88 (2015) 662673single phase liquid was not achieved for every test [26].Nonetheless, because the system experienced full vacuum condi-tions, and because parasitics were signicantly smaller here versusthe LN2 tests (see Section 4.2), liquid droplets form along the wallalmost immediately. This leads to signicantly faster chilldowntimes relative to the lower Re LN2 case.

4.2. Inner wall temperature and inner wall heat ux

To generate reverse boiling curves, the method of Burgraff [4]can be used to compute inner wall temperature from the measured

liquid hydrogen line chilldown test for TSAT = 20.3 K, 0.014 kg/s, Re = 1.22 105.

outer wall temperature. With knowledge of the parasitic heat leak,inner wall temperature can then be used to compute the inner wallheat ux. The inner wall temperature is determined from a trun-cated series solution using the assumption of adiabatic outer wall:

Ti To r2o

4a2

riro

12ln ri

ro

dTodt

164a2

r4i 5r4o r2or2i

8a2ln

riro

r

4o

16a2ln

riro

r

2or

2i

16a2

d2Todt2

9The inner and outer radii are known, as are thermal properties

of the metal and Pyrex. Outer wall temperatures from the data arefed into Eq. (9) to obtain inner wall temperature as a function oftime. Inner wall heat ux is then obtained from the Burgraffmethod (1964). Total convective heat ux is then computed by

J. Hartwig et al. / International Journal of Heausing the heat conduction equation:

q00i q00w q00axial rori

q00rad q00solidcond q00gascond

10

where the axial term is assumed to be negligible [16,17]. Table 1lists the estimated parasitic heat leaks incurred during the highRe LH2 chilldown experiments from [26] and the low Re LN2 chill-down experiments from [9], along with associated uncertainties.The cold section of the LN2 tests was the portion contained withinthe vacuum section while the warm section was the portion of theline in between liquid storage dewar and vacuum section. For LH2tests, the cold section was the horizontal portion of the transferline assembly that always contained liquid; the warm sectionwas the portion of the line initially at 250 K. The total estimatedheat leak for the entire LN2 and LH2 line chill assemblies from stor-age dewar to a point downstream of the sight glass is estimated tobe 175 W and 19W, respectively. Clearly, the total parasitic energyinto the system over the duration of the test is heavily dependenton the total chill down time. For a typical test, at the point of steadystate, the average parasitic heat leak energy into the LN2 line chillassembly was approximately 25.5 kJ while the average parasiticheat leak into the LH2 line chill assembly was 6 kJ. Uncertaintiesfor the LN2 parasitics are listed in Table 1. Uncertainties for thecomputed inner wall heat uxes for the LH2 tests are estimated in[8].

Fig. 7 plots two reverse boiling curves for the low Re LN2 exper-iments (Re = 3500) from [9] (Fig. 7(a) and (b)) versus the high ReLH2 experiments (Re = 3.11 105) from the current work(Fig. 7(c) and (d)). Fig. 7(a) and (b) show that the LN2 tests tra-versed the reverse boiling curve through vapor lm to transition

Table 1Estimated parasitic heat leaks at steady state for liquid nitrogen and liquid hydrogenchilldown tests.

Liquid nitrogen tests Liquid hydrogen tests

Heattransfer rate[W]

Uncertainty(%)

Heattransfer rate[W]

Uncertainty(%)

Warm lineradiation

2.65 11 10.91 15

Cold lineradiation

0.33 17 5.69 10

Sensors 0 0 1.91 15Solid

conduction0.05 35 0.05 40

Warm line gasconduction

172 1 0.02 15

Cold line gas 0.3 10 0.01 15

conduction

Total 175.33 10 18.59 9to nucleate boiling regimes. The maximum heat ux measured atthe sight glass was roughly 8000W/m2, and inner wall temperaturedropped to near LN2 boiling temperatures after 2 min.

Fig. 7(c) and (d) show that the high Re LH2 tests enter immedi-ately into transition boiling, and that maximum heat ux isachieved within seconds. The immediate entrance into transi-tion/nucleate boiling is shown for all three skin diode locationson the transfer line. Flow visualization during testing substantiatesthis claim, because liquid droplets were observed along the sightglass wall almost immediately after the start of the test. Inner walltemperatures follow outer wall temperatures by dropping rapidlyto LH2 saturation temperatures. At higher ow rates, the maximumheat ux through the stainless steel pipe is almost 150,000W/m2,and the maximum inner wall heat ux at the sight glass is about40,000W/m2. Heat ux is higher at upstream locations due tolower quality and a greater difference between stream and walltemperature upstream. Comparing the heat ux at SD17 andSD18 to the heat ux at SD19, heat ux at the stainless steel pipingis signicantly higher over the Pyrex test section.

The heat ux at SD19 suggests that ow at the sight glass maybe close to lm boiling at the start of the test. This correlates withthe visualization of vapor ow at the sight glass at the beginning ofeach test. The high heat uxes at SD17 and SD18 at the beginningof the test imply immediate liquid contact on the pipe upstream ofthe sight glass. Therefore higher number Re ows using LH2 areassociated with an immediate entrance into transition boiling, fas-ter chilldown, and higher maximum inner wall heat uxes relativeto low Re number LN2 ows.

4.3. Heat transfer coefcient

Heat transfer coefcients at the inner wall can now be com-puted from knowledge of the temperature difference betweeninner wall and local saturation temperature, and knowledge ofthe inner wall heat ux from Eq. (10):

hi q00i

Ti Tsat 11

A more accurate value of heat transfer coefcient could beobtained by using the stream temperature, however, the use of sat-uration temperature is widely accepted in the two phase ow com-munity. Uncertainties in the computed heat transfer coefcientsare given in [8].

Fig. 8(a) and (b) compare the heat transfer coefcient obtainedfrom the low Re number (Re = 3500) LN2 tests conducted by Huet al. [9] with that obtained from the high Re number(Re = 3.11 105) LH2 ows from the current experiment. Similarto the boiling curve plots, the LN2 heat transfer coefcient plot ischaracterized by two local inection points corresponding to theLeidenfrost point and point of maximum heat ux. Heat transfercoefcient reaches a maximum at about 200W/m2 K for low Reows; for the LH2 experiment, heat transfer coefcient was calcu-lated at SD17, SD18, and SD19. Similar to the boiling curve plots forLH2, only one maximum occurs, associated with the point of max-imum heat ux at the change from transition to nucleate boiling.Heat transfer coefcient is again higher for upstream diodes. Thisis due to the higher heat uxes associated with upstream locationsdue to lower quality ow and larger temperature gradient betweenwall and uid. This result is supported by Westbye et al. [39] and isreportedly due to greater subcooling and liquid content upstream.Heat transfer coefcient calculated at metal tube sections droppednear the boiling point of LH2 while the heat transfer coefcient atthe sight glass dropped at roughly 50 K. This behavior is attributed

t and Mass Transfer 88 (2015) 662673 669to the lower thermal conductivity of Pyrex versus SS, which is onaverage 6% of the SS value across the full temperature rangeincurred during chilldown experiments. Heat transfer coefcient

er W

Te

Hea50

100

0 50 100 150

Inn Nucleate Boiling150

200

250

300

all T

empe

ratu

re [K

]

Film Boiling

Transition Boiling

(a)

670 J. Hartwig et al. / International Journal ofacross the metal piping peaked at about 1000 W/m2 K while heattransfer coefcient across the Pyrex sight glass reached roughly700W/m2 K.

5. Discussion of results

5.1. Effect of Reynolds number

Obviously, higher Reynolds numbers will equate to higher heattransfer between a warm tube and cold cryogenic uid. Due to thehigh velocity, the LH2 test has higher Reynolds number, whichcauses more vortices in the mean ow, decreasing the thicknessof the boundary layer and further increasing the boiling heat trans-fer coefcient. Therefore, a shorter chilldown time is expected. Thehigh velocity also increases the chance of liquid droplets forming

Time [s]

0

50

100

150

200

250

300

0 20 40 60 80 100 120 140 160

SD17SD18SD19

Inn

er W

all T

empe

ratu

re [K

]

Time [s]

Transition Boiling

Nucleate Boiling

(c)

Fig. 7. (a) LN2 inner wall temperature prole, (b) LN2 reverse boiling curve, (c) LH2 inner wRe = 3.11 105.

0

50

100

150

200

0 50 100 150 200 250

Hea

t Tra

nsf

er Co

effic

ient [

W/m

2 K]

TWall

- TSat

[K]

(a)

Fig. 8. Heat transfer coefcients for (a) Low Re LN2 tests and (b) High Re50

100

150

0 1000 2000 3000 4000 5000 6000 7000 8000

Inner

W

all200

250

300

mpe

ratu

re [K

] (b)

t and Mass Transfer 88 (2015) 662673along the wall earlier in the chilldown process. At higher Re, liquiddroplets gain momentum and cause violent mixing with the vaporlm, negating the traditional inverted annular ow regime seen inmany other uids. Therefore, a higher Leidenfrost temperature isalso expected due to the high ow rate conditions.

5.2. Property effect

Comparison of the boiling curves in Fig. 7 between nitrogen andhydrogen shows that hydrogen nearly instantaneously bypassesthe lm boiling regime at high Re, with only a small fraction of timespent in transition boiling. Meanwhile, a lm boiling regime is pre-sent in all of the nitrogen data. The cause for the discrepancy inbehavior between uids is attributed to the special properties ofhydrogen. First, the liquid/vapor density ratio for hydrogen is

Inner Wall Heat Flux [W/m2]

0

50

100

150

200

250

300

0 50000 100000 150000 200000 250000

SD17SD18SD19I

nn

er W

all T

empe

ratu

re [K

]

Inner Wall Heat Flux [W/m2]

(d)

all temperature prole, and (d) LH2 reverse boiling curve. Re = 3500 for LN2 test and

0

500

1000

1500

2000

0 50 100 150 200 250

SD17SD18SD19

Hea

t Tra

nsf

er Co

effic

ient [

W/m

2 K]

TWall

- TSat

[K]

(b)

LH2 tests. ReD = 3500 for LN2 test and Re = 3.11 105 for the LH2 test.

Heasignicantly smaller than that of nitrogen. Second, the surface ten-sion of hydrogen is nearly an order of magnitude smaller than thatof nitrogen. Third, the kinematic viscosity of hydrogen is less thannitrogen. Whereas nitrogen can maintain a liquid core structuredue to large differences in properties between vapor and liquid,ow visualization from LH2 chilldown tests showed that high Rehydrogen ows did not support the inverted annular ow structure.The higher surface tension of nitrogen may act to divide the inter-faces whereas hydrogen may act like two gases mixing togetherso that transition boiling is inevitable. Therefore, to some extent,the ow boiling of hydrogen at higher Re can be regarded ashomogenous ow. Obviously, due to the limited sampling rate ofthe data acquisition system, limited resolution of the Burggrafmethod, and high Re ow, the Leidenfrost temperature cannot bediscerned in any of the LH2 chilldown data. Fourth, for the samemass ux, the ratio of Weber numbers for hydrogen over nitrogenis 53; the ratio of Froude numbers is 12. Both of thesenon-dimensional numbers have a fundamental impact on the twophase ow behavior and transition between different ow regimes.

5.3. Energy analysis

Based on the energy balance from Eqs. 18, Shaeffer et al. [31]introduced an efciency parameter to gauge the performance ofdifferent chilldown methods:

ghfg QlineQflow

QlineR tSSt0 _mhfgdt

12

The calculated efciency for the low Re trickle ow LN2 exper-iments from [9] is less than 2%, which is comparably less than thecalculated efciencies for the pulse LN2 chilldown experimentsfrom [31]. Differences in efciencies between the two LN2 experi-ments is attributed to the different range of Re numbers, due tohigher efciencies in pulse over trickle ows, and due to improve-ments in mitigating parasitic heat leak.

Based on a similar energy balance, an efciency parameterbased on actual uid inlet and exit enthalpies was created for theLH2 tests to compare performance between the different chilldownmethods, inlet liquid temperatures, and mass ow rates, and thoseresults are presented in [8]. For the 20.3 K saturated LH2 experi-ments, efciencies ranged from 0.1 to 0.56. This is a full order ofmagnitude higher efciency over the low Re number trickle owtests from [9] and mid-range Re pulse ow tests from [31].Higher efciencies for the LH2 tests over the LN2 tests are attribu-ted to the lower parasitic heat leak and higher ow rates, whichcauses turbulent mixing of the uid and faster chill down.

5.4. Critical heat ux

The critical heat ux (CHF) can also be used to compare the twoexperiments. The CHF is a point of maximum heat ux betweenwall and uid between transition and nucleate boiling. Over thepast half century, numerous correlations have been proposed forcalculating the CHF. Zuber [45] proposed the following model forstationary uid, pool boiling, which holds for when the vaporliq-uid interface of the escaping vapor passage becomes unstable dueto the Helmholtz instability:

q00CHF 0:131qvhfgclvgql qv

q2v

0:2513

This pioneering correlation is based on the two-phase hydrody-namic instability for the boiling on an innite, upward-facing, hor-

J. Hartwig et al. / International Journal ofizontal at plate. Lienhard and Dhir [22] rened Zubers model byassuming that the Helmholtz instability wavelength is equal to theTaylor instability wavelength:q00CHF 0:149qvhfgclvgql qv

q2v

0:2514

For ow boiling, the ground work was produced by Kattosgroup (19801987). The correlation format proposed by Kattosgroup is also based on the theory of the Helmholtz instabilityon the interface between vapor and liquid. They also assumedthat the CHF appears when the heat from the heated surfaceexactly balances the latent heat of the total evaporation of theliquid owing into the liquid lm layer, and the density ratiobetween vapor and liquid is sufciently smaller than unity.Katto and Kurata [12] and Katto and Ohno [13] proposed the fol-lowing functional relationship for the ow boiling CHF with noinlet subcooling:

q00CHF 0:186Glhfgqvql

0:559We0:264l 15

where We is the Weber number dened as:

Wel G2l L=qlclv 16This correlation is backed by a large set of experimental data for

multiple uids (water, refrigerants, liquid helium) as well asgeometries (tube, at plate) [14,15]. Mudawar and Maddox [25]further modied the CHF correlation to explicitly handle geometri-cal effects:

q00CHF 0:161Glhfgqvql

1523

We 823l

Ld

123

17

The above four correlations were applied to the low Re numberLN2 tests from [9] and the high Re LH2 tests from [8]. Table 2 listsresults. For the LN2 tests, correlations were applied to the test sec-tion inside the vacuum section; for the LH2 tests, correlations wereapplied to the point just upstream of the sight glass, at P3 in Fig. 4for the 24.2 K saturated high ow case.

Regarding the LN2 tests, all four correlations over-predict CHFvalues. One reason for this is attributed to the low level of vac-uum leading to relatively large heat leakage. The heat leakagewill introduce some vapor into the bulk ow which leads to arelatively high quality saturated ow boiling rather than sub-cooled ow boiling. Furthermore, the increasing quality acceler-ates the bulk ow due to the buoyancy force, which increasesthe ability to sustain a continuous liquid ow on the tube wall.The phenomenon is also reinforced by the pressure oscillationduring the intensive boiling. Discrepancies between measuredand computed CHF may also be attributed to thermal hysteresis;the CHF is lower during quenching than heating on a givensurface.

Compared to the measured heat ux for LH2 tests, the Zuber[45] and Lienhard and Dhir [22] correlations underpredict andthe Katto group over predicts the data. (19801987). The Zuberand Lienhard and Dhir correlations are for pool boiling whereasexperiments here are for ow boiling where it is expected to havehigher heat transfer rates with the strong vortex convection.Kattos correlation was derived assuming external ow over aheated plate whereas experiments here were for internal pipe ow.Meanwhile the Mudawar and Maddox [25] correlation was vali-dated for 0:0095 < qvql < 0:0102; density ratio for LH2 tests was

0.04, much higher than the validation range. Another cause for dis-crepancy between Mudawar and Maddox and the LH2 data isattributed to the assumption that the two-phase ow must havea steady vapor-blanket for the correlation to be valid. Flow visual-

t and Mass Transfer 88 (2015) 662673 671ization from Fig. 6 shows that annular ow is achieved relativelyquickly, and that the system spends little time in vapor lmcooling.

Measured maximum heat ux 13.8 233.4Model

Hea5.5. Leidenfrost point

Another critical parameter for the chilldown process is theLeidenfrost temperature, or rewet temperature. Based on theTaylor wave action, Berenson [1] derived the following formulato calculate the Leidenfrost temperature:

Tlei;B Tsat DT

0:127qvhfgkv

gql qvlvql qv2

!0:5g0clv

gql qv 0:33

18

Carey [5] developed a model of Leidenfrost phenomenon basedon liquid microlayer evaporation:

s " #0:6

Zuber [45] 162 87.2Lienhard and Dhir [22] 182 99.2Kattos group (19801987) 92 377.4Mudawar and Maddox [25] 24.3 133.9

Table 3Comparison of measured vs. calculated Leidenfrost temperatures for the LN2 and LH2chilldown experiments.

LN2 experiment LH2 experimentLeidenfrost T [K] Leidenfrost T [K]

170.2 ModelBerenson [1] 143.8 60.8Carey [5] 144.5 287Table 2Comparison of measured vs. calculated critical heat ux (CHF) for the LN2 and LH2chilldown experiments.

LN2 experiment LH2 experimentHeat ux [kW/m2] Heat ux [kW/m2]

672 J. Hartwig et al. / International Journal ofTlei;H Tlei;BTlei;B Tl 0:42

klqlcplkwqwcpw

hfgcpwTlei;B Tsat 19

Leidenfrost temperatures are computed using these two modelsand compared with the actual experimental value; results are pre-sented in Table 3. Note from Fig. 7(c) and (d) that the LH2 testsnever passed through the Leidenfrost point, so there is no experi-mental data to compare with the theory. Both correlations predictreasonably well with the LN2 measured value. Meanwhile,Berensons [1] correlation predicts a low Leidenfrost temperaturefor LH2 tests. Differences between the correlations are attributedto the fact that Berensons correlation was determined for transi-tion pool boiling over a horizontal surface whereas the data wastaken for internal, forced two-phase ow. Meanwhile, Careys [5]correlation interestingly predicts a Leidenfrost temperature abovethe warmest initial temperature of 250 K for the LH2 tests. Becausethe vacuum chamber wall for the LH2 tests was set to 250 K ambi-ent, there is no way to substantiate or negate this correlation.

6. Conclusion

A systematic comparison has been made between two differentchilldown experiments conducted in two different uids havingdifferent ow regimes. Analysis of the temperature traces, owvisualization, heat ux, parasitic heat leaks, heat transfer coef-cients, differences in properties, critical heat ux, and Leidenfrostpoints indicates stark differences in chilldown behavior between

hydrogen transfer line chilldown experiments. II. Analysis, Int. J. Multiphase

chilldown of a simulated exible metal hose using liquid nitrogen, J. LowTemp. Phys. 174 (2014) 247268.[11] T. Jin, J. Hong, H. Zheng, K. Tang, Z. Gan, Measurement of boiling heat transfercoefcient in liquid nitrogen by inverse heat conduction method, J. ZhejiangUniv. Sci. A 10 (2009) 691696.

[12] Y. Katto, C. Kurata, Critical heat ux of saturated convective boiling onuniformly heated plates in a parallel ow, Int. J. Multiphase Flow 6 (1980)575582.

[13] Y. Katto, H. Ohno, An improved version of the generalized correlation of criticalheat ux for the forced convective boiling in uniformly heated vertical tubes,Int. J. Heat Mass Transfer 27 (1984) 16411648.

[14] Y. Katto, S. Yokoya, Critical heat ux of liquid helium (I) in forced convectiveboiling, Int. J. Multiphase Flow 10 (1984) 401413.

[15] Y. Katto, S. Yokoya, Critical heat ux of forced convective boiling in uniformlyheated vertical tubes with special reference to very large length-to-diameterratios, Int. J. Heat Mass Transfer 30 (1987) 22612269.

[16] O. Kawanami, H. Azuma, H. Ohta, Effect of gravity on cryogenic boiling heattransfer during tube quenching, Int. J. Heat Mass Transfer 50 (2007) 34903497.

[17] O. Kawanami, T. Nishida, I. Honda, Y. Kawashima, H. Ohta, Flow and heattransfer on cryogenic ow boiling during tube quenching under upward anddownward ow, Microgr. Sci. Technol. 19 (34) (2007) 137138.

[18] S. Kim, I. Mudawar, Universal approach to predicting heat transfer coefcientfor condensing mini/micro-channel ow, Int. J. Heat Mass Transfer 56 (2013)238250.Flow, (in press).[9] H. Hu, J.N. Chung, S.H. Amber, An experimental study on ow patterns and

heat transfer characteristics during cryogenic chilldown in a vertical pipe,Cryogenics 52 (2012) 268277.

[10] H. Hu, T.K. Wijeratne, J.N. Chung, Two-phase ow and heat transfer duringthe two uids. Analysis and discussions in the current work showthat the chilldown process evolves much more rapidly in hydrogenover nitrogen due to lower density ratio, surface tension, kinematicviscosity, and due to the quick transition from vapor to annularow. Meanwhile lm boiling is much more pronounced in nitrogenchilldown tests due to the higher density ratio, higher surface ten-sion, and higher kinematic viscosity. Second, higher ow rate LH2chilldown tests yield higher efciencies than the correspondinglower ow rate LN2 experiments; discrepancies are primarilyattributed to differences in parasitic heat leak. Third, the criticalheat ux correlations used to compute the maximum heat uxare shown to over-predict the LN2 data, but generallyunder-predict the LH2 data. Lastly, correlations for predictingLeidenfrost temperatures agree reasonably well with LN2 databut cannot be used to compare with the LH2 data. Results herehave direct implications on the design of all future cryogenic pro-pellant transfer systems.

Conict of interest

None declared.

References

[1] P.J. Berenson, Transition Boiling Heat Transfer from a Horizontal Surface,Technical Report 17, Massachusetts Institute of Technology, March 1, 1960.

[2] L. Bewilogua, R. Knoner, H. Vinzelberg, Heat transfer in cryogenic liquids underpressure, Cryogenics 15 (1975) 121125.

[3] J.A. Brennan, E.G. Brentari, R.V. Smith, W.G. Steward, An Experimental Reporton Cooldown of Cryogenic Transfer Lines, NBS Report 9264, Boulder CO,November 7, 1966.

[4] O.R. Burgraff, An exact solution of the inverse problem in heat conductiontheory and applications, ASME J. Heat Transfer 86 (1964) 373380.

[5] V.P. Carey, Liquid Vapor Phase Change Phenomena: An Introduction to theThermophysics of Vaporization and Condensation Processed in Heat TransferEquipment, second ed., Taylor and Francis, New York, 2008.

[6] U.H. Glahn, A Correlation of Film-Boiling Heat-Transfer Coefcients Obtainedwith Hydrogen, Nitrogen, and Freon 113 in Forced Flow, NASA-TN-D-2294,May 1, 1964.

[7] RW. Graham, R.C. Hendricks, R.C. Ehlers, An experimental study of the poolheating of liquid hydrogen in the subcritical and supercritical pressure regimesover a range of accelerations, Adv. Cryog. Eng. 10 (1965) 342352.

[8] J.W. Hartwig, J. Styborski, B. Stiegemeier, H. Hu, E. Rame, J.N. Chung, Liquid

t and Mass Transfer 88 (2015) 662673[19] S. Kim, I. Mudawar, Universal approach to predicting saturated ow boilingheat transfer in mini/micro-channels part I. Dryout incipience quality, Int. J.Heat Mass Transfer 64 (2013) 12261238.

[20] S. Kim, I. Mudawar, Universal approach to predicting saturated ow boilingheat transfer in mini/micro-channels part II. Two-phase heat transfercoefcient, Int. J. Heat Mass Transfer 64 (2013) 12391256.

[21] B. Kutter, F. Zegler, G. ONeil, B. Pitchford, A Practical, Affordable CryogenicPropellant Depot Based on ULAs Flight Experience, AIAA-2008-7644, 2008SPACE Conference, San Diego, CA, September 911, 2008.

[22] J.H. Lienhard, V.K. Dhir, Hydrodynamic prediction of peak pool boiling heatuxes from nite bodies, J. Heat Transfer 95 (1973) 152158.

[23] S.M. Liao, T.S. Zhao, An experimental investigation of convection heat transferof supercritical carbon dioxide in miniature tubes, Int. J. Heat Mass Transfer 45(2002) 50255034.

[24] B. Louie, W.G. Steward, Onset of nucleate and lm boiling resulting fromtransient heat transfer to liquid hydrogen, Adv. Cryog. Eng. 35 (1990) 403412.

[25] I. Mudawar, D.E. Maddox, Enhancement of critical heat ux from high powermicroelectric heat sources in a ow channel, J. Electron. Packag. 112 (1990)241248.

[26] E. Rame, J.W. Hartwig, J.B. McQuillen, Flow Visualization of Liquid HydrogenLine Chill Down Tests, AIAA-2014-1074, 2014 AIAA SciTech Conference,National Harbor, MD, January 1317, 2014.

[27] W.M. Rohsenow, A Method of Correlating Heat Transfer Data for SurfaceBoiling of Liquids, Technical Report #5, Ofce of Naval Research, July 1, 1951.

[28] J.R. Schuster, D.J. Howell, S.L. Lucas Jr., M.S. Haberbusch, J.D. Gaby, N.T. VanDresar, M.F. Wadel, Cold Flow Testing of Revised Engine Chilldown Methodsfor the Atlas Centaur, 32nd AIAA-96-3014, 32nd Joint Propulsion Conference,Lake Buena Vista, FL, July 13, 1996.

[29] J.R. Schuster, C.E. Bassett, E.H. Christensen, S.C. Honkonen, F. Merino, D.S.Munko, J.R. Pietrzyk, M.A. Wollen, Centaur-Based Liquid Hydrogen FluidManagement Flight Experiments, AIAA-91-3539, AIAA/NASA/OAI Conferenceon Advanced SEI Technologies, Cleveland, OH, September 46, 1997.

[30] J.D. Seader, W.S. Miller, L.A. Kalvinskas, Boiling Heat Transfer for Cryogenics,NASA-CR-243, June 1965.

[31] R. Shaeffer, H. Hu, J.N. Chung, An experimental study on liquid nitrogen pipechilldown and heat transfer with pulse ows, Int. J. Heat Mass Transfer 67(2013) 955966.

[32] J.E. Sherley, Nucleate boiling heat-transfer data for liquid hydrogen at standardand zero gravity, Adv. Cryog. Eng. 8 (1963) 495500.

[33] Y. Shirai, H. Tatsumoto, K. Hata, M. Shiotsu, Preliminary study on heat transfercharacteristics of liquid hydrogen for coolant of HTC superconductors, Adv.Cryog. Eng. 55 (2010) 337344.

[34] Y. Shirai, H. Tatsumoto, M. Shiotsu, K. Hata, H. Kobayashi, Y. Naruo, Y. Inatani,K. Kinoshita, Forced ow boiling heat transfer of liquid hydrogen forsuperconductor cooling, Cryogenics 51 (2011) 295299.

[35] J.L. Sloop, Liquid Hydrogen as a Propulsion Fuel, 19451959, NASA-SP-4404,1978.

[36] H. Tatsumoto, Y. Shirai, K. Hata, T. Kato, M. Shiotsu, Forced convection heattransfer of subcooled liquid nitrogen in a horizontal tube, Adv. Cryog. Eng. 47B(2008) 665672.

[37] H. Tatsumoto, Y. Shirai, K. Hata, T. Kato, M. Shiotsu, Forced convection heattransfer of nitrogen at supercritical pressure, in: 2008 International CryogenicEngineering Conference Seoul, Korea, 2008, pp. 383388.

[38] C. Tanner, J. Young, R. Thompson, A. Wilhite, On-Orbit Propellant ResupplyOptions for Mars Exploration Architectures, AIAA-2006-261, IAC-06-D1.1.01,2006.

[39] C.J. Westbye, M. Kawaji, B.N. Antar, Boiling heat transfer in the quenching of ahot tube under microgravity, AIAA J. Thermophys. Heat Transfer 9 (1995) 302307.

[40] K. Yuan, Cryogenic Boiling and Two-Phase Chilldown Process underTerrestrial and Microgravity Conditions (Ph.D. thesis), University of Florida,2006.

[41] K. Yuan, Y. Ji, J.N. Chung, Cryogenic chilldown process under low ow rates, Int.J. Heat Mass Transfer 50 (2007) 40114022.

[42] K. Yuan, Y. Ji, J.N. Chung, W. Shyy, Cryogenic boiling and two-phase owduring pipe chilldown in earth and reduced gravity, J. Low Temp. Phys. 150(2008) 101122.

[43] K. Yuan, Y. Ji, J.N. Chung, Numerical modeling of cryogenic chilldownprocess in terrestrial gravity and microgravity, Int. J. Heat Fluid Flow 30(2009) 4453.

[44] P. Zhang, X. Fu, Two-phase ow characteristics of liquid nitrogen in verticallyupward 0.5 and 1.0 mm micro-tubes: visualization studies, Cryogenics 49(2009) 565575.

[45] N. Zuber, Hydrodynamic aspects of boiling heat transfer, AEC Report No. AECU-4439, 1959.

J. Hartwig et al. / International Journal of Heat and Mass Transfer 88 (2015) 662673 673

Comparison of cryogenic flow boiling in liquid nitrogen and liquid hydrogen chilldown experiments1 Introduction2 Background and governing physics3 Experimental setup3.1 LN2 experiments3.2 LH2 experiments

4 Results and comparison4.1 Temperature traces and flow visualization4.2 Inner wall temperature and inner wall heat flux4.3 Heat transfer coefficient

5 Discussion of results5.1 Effect of Reynolds number5.2 Property effect5.3 Energy analysis5.4 Critical heat flux5.5 Leidenfrost point

6 ConclusionConflict of interestReferences