Direct Contact Heat Transfer

8/12/2019 Direct Contact Heat Transfer

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INTRODUCTION

There are many process applications of gas/liquid directcontact equipment where the main objective is transfer ofheat. The preference for this type of equipment is due to itsrather simple design (leading to reduced capital expenses)and ease of operation since heat transfer surfaces (and theirusual corrosion and fouling problems) are absent, thussignificantly reducing operating and maintenance expenses.

Considerable work has been carried out on packed bedtype direct-contact condensers (e.g. Fair, [1]). However,well-established methods for predicting heat transfer rates insuch devices are still unavailable. The situation is worse in

the case of structured packed beds, which are very attractivefor a variety of applications. Indeed, there is a lack of heattransfer data in the open literature, helpful in clarifying somekey issues involved and in providing guidance for designcalculations [2], [3].

A complication frequently arising in studies of direct-contact heat transfer is due to the simultaneous transfer ofmass which may be in the same direction as the main heatflow or may proceed opposite to it, depending on processconditions. Therefore, experimental determination of overallheat transfer coefficients over the entire packed column maymask such effects and may not help improve ourunderstanding of this process. In one of the few availabledirect-contact heat transfer studies on structured packings,Spigel et al [4] obtained data with air/water and an air/oilsystem. Overall transfer units and overall heat transfercoefficients U were reported for packed columns 0.63 and0.675 m high. Correlations were also proposed for

predicting heat transfer rates for structured packings. For thecase of sub-cooled water brought in contact with hot air,they reported that U varied with the gas Reynolds number toa power 0.8. Huang and Fair [5] have obtained aconsiderable amount of data on the air/water system invarious types of packed columns but not in structuredpackings. Bontozoglou and Karabelas [6] carried outexperiments with the same structured packing employed inthis work utilizing steam, with a small percentage of CO2,flowing upwards in a column where sub-cooled water wasfed at the top. They reported that the condensationcoefficient was enhanced at the lower levels of the column(close to the steam feeding point), which was attributed to

the fact that during the early stages of gas/liquid contact, thegas (mainly saturated steam) is poor in non-condensables.However, as the distance from the feeding point increases(and so does the gas/liquid contact time) more steamcondenses on the gas/liquid interface and the mixturebecomes richer in non-condensables which retardcondensation.

The scope of this work is to collect reliableexperimental data in a pilot scale column with structuredpacking that would enrich the literature and would offer theopportunity to improve understanding of the direct-contactprocess along the column. The common air/water system isemployed which (in addition to its practical usefulness)

allows comparison with a few relevant literature studies.

DIRECT CONTACT AIR-WATER HEAT TRANSFER

IN A COLUMN WITH STRUCTURED PACKING

Sofoklis Kypritzis and Anastasios J. Karabelas

Department of Chemical Engineering and Chemical Process Engineering Research Institute,Aristotle University of Thessaloniki, Univ. Box 455, GR 540 06 Thessaloniki, Greece

Tel: (+30) 31 996201, Fax: (+30) 31 996209, E-mail: [email protected]

ABSTRACT

This paper deals with the process of direct contact heat transfer between hot (relatively dry) air and sub-cooledwater flowing counter-currently in a column filled with structured packing (Sulzer, Melapak 250.Y). The mainobjective is to collect reliable experimental data for a type of process equipment (structured packed column)inadequately studied so far. The range of flow parameters studied was as follows: Air flux between 2.4 and 6.2tn/hm

2 and Reynolds number between 820 and 2150; water flux between 7.2 and 17.2 tn/hm

2 and Reynolds

number between 10.4 and 25. The pressure varied from atmospheric to 1 bar gauge. A fully instrumented pilotscale unit was employed in the tests with a stainless steel column 70cm long and 15cm i.d. A significant amountof new data has been collected by measuring the local temperature at various levels of the column.

Sensible heat exchange between air and water in the column takes place simultaneously with mass transfer.Interpretation of data shows that the process strongly depends on the air flow rate, as expected. The influence ofliquid rate is noticed only in the upper section of the column. There is evidence that a very significant amount ofheat, exchanged between the two streams is due to latent heat (mass) transfer. The bottom part of the packed bedoperates essentially as an air cooling section promoting saturation of the supplied hot air. Under some conditions,both air humidification and subsequent de-humidification may take place in the bottom section i.e. within a rathershort packing segment. Above this section, the device seems to operate as an ordinary direct contact heater of theliquid phase. For the middle and upper sections, common correlations provide predictions in rough agreementwith measured heat transfer coefficients.

Session Thermalhydraulics: Fluidizedand Packed Beds 1

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EXPERIMENTAL EQUIPMENT AND PROCEDURE

The experimental system is shown in Figure 1. Thecondenser is a 1,05 m long, 0,15 m I.D. column made ofstainless steel 316. A hot well, equipped with a magneticfloater forms the bottom of the column. The liquid outletvalves are electrically actuated through the floater in order to

prevent the passage of air in the drainage pipe. The columnis filled with Mellapak 250.Y structured packing marketedby Sulzer, which appears to be one of the packing typesfavored for direct contact applications. Characteristics ofthis packing are listed in Table 1.

Tap water is demineralized and delivered to a spraymanifold at the top of the column. The water flow rate iscontrolled by a PID controller and a flowmeter whichelectronically actuate a Badger control valve. Dry air isprovided by the Laboratory compressor facility. Heatingtakes place in a special vessel equipped with electricalresistances (3,5 KW). Pressure gauges and flowmeters areinstalled to monitor air flow conditions. The air heatingvessel, piping and packed column are thermally insulated toavoid heat losses and achieve adiabatic conditions. Atemperature controller is installed at the outlet of the airheating vessel to maintain constant temperature (120

oC)Temperatures are monitored with K-type thermocouples

calibrated to 0,2oC. Thermocouples are installed in the well(TI1), at the cold water inlet (TI7) and at the air outlet (TI8).Five thermocouples are embedded in the packing (identifiedby TI2, TI3, TI4, TI5, TI6, in Figure 1) located at a distance80, 120, 160, 375, 585 mm (respectively) from the bottom ofthe packing. The thermocouple tips are positioned close tothe centerline of the column, in contact with the packing toprovide the temperature of the local liquid film.Temperatures are indicated on a central panel and are also

recorded by a data acquisition system for later processing.The range of conditions tested is as follows:

Air flux between 2.4 and 6.2 tn/hm2 and Reynoldsnumber (based on the effective velocity of air) between 820and 2150.

FIG. 1: The experimental device

Water flux between 7.2 and 17.2 tn/hm2 and Reynoldsnumber (based on superficial water velocity) between 10.4and 25.

Pressure from atmospheric up to 1 bar gauge.To facilitate data interpretation the packed bed is consideredto be comprised of five sections henceforth to be referred to

as top, upper, middle, lower and saturation sections, asindicated in Figure 1. The locations of thermocouples TI3,TI4, TI5 and TI6 serve as the boundaries of these sections.Thermocouple TI2 provides an estimate of saturation sectiontemperature.

RESULTS

Water Temperature Profiles

Typical water temperature profiles measured in these testsare plotted in Figures 2,3 and 4. These data show that wateris heated (as expected) while flowing through the top, upper,middle and lower sections. A modest water heating is

achieved in the three top sections, possibly due to sensibleheat transfer together with some humidity (mass) transportfrom the saturated warm air to the falling water films, underconditions of forced convection. A rather sharp temperatureincrease takes place (up to a maximum) within the lowersection. At first look, it is doubtful whether convective heattransfer is the main mechanism responsible for thissignificant heating. Of greater interest is the reduction ofwater temperature within the so-called saturation section.This is apparently caused by the humidification of theincoming hot dry air, effected through evaporation andleading to water cooling. This explanation is supported bythe observation that thermocouples TI1 and TI2

systematically measure water temperatures lower than TI3,as shown in figures 2,3,4. In summary, it appears that hot air

Cold water inlet

Cold & humid air outlet to

Atmosphere.

Hot & dry air inlet,

Heated water outlet

to drain

TIC

TI8

FIC

PI

FI

PI

LC

TI7

Tap water inlet

Air heatingdevice

Pressurized dryair at 8 bar abs

PITI2

TI3

TI4

TI5

TI6

TI1

700mm

Lower section

Middle section

Saturation

section

Upper section

Top section

Waterdemineralization

2001, ExHFT-5, Thessaloniki, Greece

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10

15

20

25

30

35

0 200 400 600 800

Distance from the bottom of the packing (mm)

Watertemperature(C)

P = 2.0 bar, Re(g) = 1320

P = 1.5 bar, Re(g) = 1340

P = 1.25 bar, Re(g) = 1280

P = 1.0 bar, Re(g) = 1380

at well

10

15

20

25

30

0 200 400 600 800

Distance from the bottom of the packing (mm)

W

atertemperature(C)

P = 1.5 bar, Re(g) = 2150

P = 1.25 bar, Re(g) = 2100

P = 1.0 bar, Re(g) = 2030

at well

10

20

30

40

50

60

70

0 200 400 600 800

Distance from the botto m of t he packed bed (mm)

Watertemperature(C)

T(wtr) P = 2.0 bar, Re(g) = 1600

T(air) P = 2.0 bar, Re(g) = 1600

T(wtr), P = 1.25 bar, Re(g) = 1590

T(air), P = 1.25 bar, Re(g) = 1590

at well

humidification of the saturated up-coming air by the colderfalling water films which (aided by sensible air/waterconvective heat transfer) leads to increasing watertemperature.

FIG. 2: Typical water temperature profiles at constant waterflux (12.2 tn/hm2) and Regas~constant

FIG. 3: Typical water temperature profiles at constant water

flux (12.2 tn/hm2) and Regas~constant

FIG. 4: Typical water and air temperature profiles at constant waterflux (12.2 tn/hm2) and Regas~constant

Air Temperature Profiles

To further interpret the data, it is essential to determinethe prevailing air temperature profiles in the column undervarious conditions. In Figure 4 calculated air temperaturesare included which were obtained, assuming adiabaticcolumn operation and saturated gaseous phase, via the

energy balance for water:

Q = Lcwtr(Tout,wtr Tin,wtr) (1)

and for humid air:

Q = G[cair(Tin,air Tout, air) + (Hin Hout)] (2)

In equations [1] and [2] all physical quantities are known;L, Tout, wtr , Tin, wtr, G and Tout, air are measured and Houtcanbe obtained via a psychrometric chart. Equation [2] may berearranged to solve for Tin, air for the corresponding Hin. Thisprocedure is repeated sequentially for all sections (exceptthe saturation section), starting from the top section whereboundary conditions (Tout, air and Tin, wtr ) are obtained fromthe corresponding thermocouples (TI8 and TI7).

Heat Transfer Coefficients

Figures 5 through 10 depict calculated overall heattransfer coefficients versus gas Reynolds number for upper,middle and bottom sections. Various water fluxes in a range7.2 to 17.22 m3/m2h are employed, for air pressure 2.0 and1.5 bar absolute. The overall heat transfer coefficient U, isobtained via equation [3] since all relevant quantities arecalculated from equations [1] and [2]. In equation [3] thelogarithmic mean temperature difference is approximated by

the arithmetic mean difference, without introducing asignificant error:

( ) ( )mlm TazAQ

TazA

QU

= [3]

Neither enhancement of interface compared to geometricpacking area, as suggested by Henriques de Brito et al [7],nor reduction due to liquid maldistribution effects isconsidered here in order to specify the effective area A forheat/mass transfer between humid air and water. Thus, thetotal effective interface area in the computations wasassumed to be constant and equal to the nominal area of thepacking (indicated in Table 1).

Figures 5 to 10 show that the heat transfer coefficient (forall three column sections considered) is strongly affected bygas Reynolds number, as expected. It is also observed that,for the lower and middle sections (Figures 6, 7, 9, 10), theheat transfer coefficient U is almost independent of theliquid flow rate. Indeed, a relatively small variation of U forvarious liquid rates (for a fixed ReG) is not systematic andappears to be within the experimental error ( 15%). Thisbehaviour of U may be attributed to a controlling resistanceto heat transfer residing at the gas side of the interface.Furthermore, the insignificant effect of liquid rate in thesesections suggests that the condition of the falling liquid filmsdoes not influence the transfer rates at the gas side.

Upon inspection of Figures 7 and 10 (for the lowersection), corresponding to absolute pressure 2 and 1.5 barrespectively, one can make the following observations:


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The magnitude of U is much greater than that for theother two sections.

Figure 10 for P=1.5 bar depicts (at high ReG ) adecreasing U with increasing ReG.

The latter trend appears to be physically impossibleand may result from violation of assumptions involved indetermining the transfer coefficient U; i.e., that only

sensible heat transfer and condensation from a saturatedgas stream take place in this section. In reality, however,it is possible that (at sufficiently high air velocities) thepacking section designated as saturation section

FIG. 5: Heat transfer coefficient vs Re of the upper section of thecolumn (P = 2.0 bar abs)

FIG. 6: Heat transfer coefficient vs Re of the middle section of thecolumn (P = 2.0 bar abs)

FIG. 7: Heat transfer coefficient vs Re of the lower section of thecolumn (P = 2.0 bar abs)

(between TI2 and TI3, Figure 1) may not be thickenough, and the air residence time there may be too short,to attain saturation. Thus, air saturation may be completedin the next section (designated as lower in Figure 1).

Furthermore, it is not unlikely that in the same (lower) section water evaporation (air humidification) may take

place first followed by condensation. Under theseconditions, Equ (2) would underestimate Q for that sectionleading to reduced U values. Apparently, with increasingReG, U would keep decreasing as shown in Figure 10.

FIG. 8: Heat transfer coefficient vs Reynolds of the upper sectionof the column (P = 1.5 bar abs)

FIG. 9: Heat transfer coefficient vs Re of the middle section of thecolumn (P = 1.5 bar abs)

FIG. 10: Heat transfer coefficient vs Re of the lower section of thecolumn (P = 1.5 bar abs)

0

20

40

60

80

100

120

0 500 1000 1500 2000 2500

Gas Reynolds number

Heattransfercoefficient(W/mxm

17,22

15,58

13,93

12,22

10,6

8,9

7,2

theoreticalprediction

0

20

40

60

80

100

120

140

0 500 1000 1500 2000 2500

Gas Reynolds number

Heattransfercoefficient

(W/mx

17,22

15,58

13,93

12,22

10,6

8,9

7,2

thoreticalprediction

0

50

100

150

200

250

300

350

400

450

0 500 1000 1500 2000 2500

Gas Reynolds number

Heattransfercoefficient(W/mxmK

17,22

15,58

13,93

12,22

10,6

8,9

7,2

0

5

10

15

20

25

30

35

40

45

0 200 400 600 800 1000 1200 1400 1600 1800

Gas Reynolds number

Heattransfercoefficient(W/mxmK)

17,22

15,58

13,93

12,22

10,6

8,9

7,2

0

20

40

60

80

100

120

0 500 1000 1500 2000

Gas Reynolds number


17,22

15,58

13,93

12,22

10,6

8,9

7,2

0

50

100

150

200

250

300

350

400

450

500

0 500 1000 1500 2000

Gas Reynolds number

Heattransfercoefficient(W/mxm

17,22

15,58

13,93

12,22

10,6

8,9

7,2


4


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The results in Figure 7 (P=2.0 bar absolute) show that(unlike those in Figure 10) U increases monotonically withReG. In line with the above explanation, this may beattributed to the higher air pressure and the relatively smallerquantity of vapour required for saturation, which may thenbe completed within the saturation section. Consequently,Equation (2) would be satisfied in the next (lower) section

with the expected U versus ReGvariation.The high U values in the lower section (e.g. Figure 7) are

difficult to explain. One may attribute this to supersaturationof air entering that section, which would lead to an amountof latent heat released greater than that accounted for in theprocedure used for determining U (equ 1 and 2).

Figures 5 and 8 as well as Figures 11 and 12 show thatwhereas for the lower and middle sections U is practicallyindependent of water flow rate, for the upper section U isroughly inversely proportional to ReL . This rather curioustrend of transfer coefficient U is not observed for the firsttime in direct-contact experiments. Bontozoglou andKarabelas [6] studying steam condensation in the same

experimental setup, reported similar trends. Karapantsios etal [8] obtained experimental data of direct-contact steamcondensation on falling water films, inside a vertical tube, inthe presence of large amounts of non condensable gases; itwas also observed that, by increasing the liquid flow rate,reduced values of integral heat transfer coefficient wereobtained. They hypothesized that (with increasing liquidrate) the enhanced liquid waves trapped noncondensablegases at the interface (or they caused a steam deplection ofthe boundary layer at the gas side) promoting a reduction ofthe mass transfer coefficient. It is uncertain whether theabove arguments apply to the present system, and this matterrequires more attention in future studies.

In view of the above observations that the resistance toheat transfer resides at the gas side, an attempt is made topredict integral heat transfer coefficients U in a section byemploying common correlations. Considering sensible heattransfer, and latent heat transported to liquid films throughturbulent convective mass transfer, one obtains U as follows:

( ) ( ) ( )

,,,,

,,,,

hTT

HHKU

HHazAKTTazAhTTazAU

QQQ

ave

wtr

ave

air

ave

satwtrT

ave

satairT

ave

satwtrT

ave

satairT

ave

wtr

ave

air

ave

wtr

ave

air

latsens

+

=

+=

+=

Following Fair & Bravo [9], Schpigel & Meier [10] one mayemploy the same type of correlation for estimating transfercoefficients K and h.

33.08.0

33.08.0

Re034.0

PrRe034.0

gasgasgas

gasgasgas

ScSh

Nu

=

=

Predictions based on this approach are plotted in Figure 8and 9 for the upper and middle sections where (as alreadydiscussed) the assumptions involved in estimating U may be

better satisfied than in other sections. Predictions are inrough (order of magnitude) agreement. It is interesting thatthe exponent of the ReGdependence of U data is closer to1.0 than to 0.8 employed in the above correlations.

FIG. 11. Heat transfer coefficient vs Re(liq)

(P =2.0 bar abs & Re(g) = 1600)

FIG. 12: Heat transfer coefficient vs Re(liq)

(P = 1.25 bar abs & Re(g) = 1470)

CONCLUDING REMARKS

The well-known air/water system employed here ischaracterized by a relatively large enthalpy ofvaporization.Therefore, if there is a large difference of inlettemperatures of the two streams, brought into direct contact,latent heat effects dominate in certain sections of the

column. Moreover, if the incoming air is dry, bothhumidification and de-humidification may take take placewithin rather short sections of the packing that requirecareful attention for data interpretation.

Under the conditions studied here, the liquid flow rateappears to have no effect on the performance of thestructured packed bed as direct-contact condenser, except atthe top of the column. In intermediate sections, where latentheat transfer is not excessive, the integral heat transfercoefficient displays a dependence on ReG to a power nearunity. In these sections, common convective mass/heattransfer correlations lead to acceptable (order of magnitude)predictions. A rather curious apparent effect (also observedin previous studies) of decreasing U with increasing liquidrate, at the top section, requires additional work to beconfirmed and clarified.

(5a)

(5b)

10

100

1000

0 5 10 15 20 25 30

Liquid Reynolds number


upper section middle section lower section

10

100

1000

0 5 10 15 20 25 30

Reynolds number of liquid


upper section middle section lower section

(4)


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SYMBOLS

A: Cross-sectional area of the column [=] m2cwtr: Specific heat of water [=] J/Kg

oC

cair: Specific heat of air [=] J/KgoC

D: Diffusivity [=] m2/s

dh: Hydraulic diameter of the packing (=4/) [=] mG: Air flow rate [=] Kg/sHin: Kg of water/Kg of dry air at the inletHout: Kg of water/Kg of dry air at the outlet

ave

sat,air,TH : Air saturation humidity atave

airT ave

sat,wtr,TH : Air saturation humidity atave

wtrT

h: Heat transfer coefficient [=] W/m2

K: Mass transfer coefficient [=] mol/m2sKg: Mass transfer coefficient [=] mol/m

2sbar

L: Water flow rate, [=] Kg/sQ: Rate of heat transferred , [=] J/sQsens: Rate of sensible heat transferred [=] J/sQlat: Rate of latent heat transferred [=] J/s

Regas: Gas Reynolds number : hogas

airgas d

AG

45cos

/Re

=

Reliq: Liquid Reynolds number :

liqpack

liqliq

liqa

u

sup,Re =

Tout,wtr: Outlet temperature of water [=]oC

Tin,wtr: Inlet Temperature of water [=]oC

Tout,air: Outlet temperature of air [=]oC

Tin,air: Inlet Temperature of air [=]oC

2

air,outair,inave

air

TTT

+=

2

wtr,outwtr,inave

wtr

TTT

+=

uliq,sup: Superficial liquid velocity [=] m/sU: Heat transfer coefficient [=] W/m2Kz: Height of packed section [=] m

Greek Letters

: Specific area of the packing (=250) [=] m-1: Latent heat of water evaporation [=] J/Kgliq/gas: Liquid/Gas viscosity respectively [=] Kg/ms

liq/gas: Liquid/Gas density respectively [=] Kg/m3

Experimental Heat Transfer, Fluid Mechanics, andThermodynamics 2001, Proceedings p.p. 1695-1700.G.P. Celata, P.Di Marco, A. Goulas and A. Mariani

2001 Edizioni ETS, Pisa. All rights reserved

REFERENCES

1. Fair J.R., Direct Contact Gas-Liquid Heat Exchangefor Energy Recovery, Trans. of ASME, Journal ofSolar Energy Engineering, vol. 112, (1990), pp. 216222

2. Fair J.R., Designing DirectContact Coolers /Condensers, Chem. Eng., Vol. 12, (1972), pp. 91 100, June

3. Fair J.R., Process Heat Transfer by Direct Fluid-PhaseContact,AIChE Symp. Ser. No. 118, vol. 68, (1971)

4. Spigel L., Bomio P., Hunkeler R., Direct heat and masstransfer in structured packings, Chem. Eng. & Proc.,35, (1996), pp. 479 485

5. Huang Chen-Chia & Fair J.R., Direct Contact Gas-Liquid Heat Transfer in Packed Column, HeatTransfer Engineering, vol. 20, no 2, (1989), pp. 19 28

6. Bontozoglou V. & Karabelas A.J., Direct ContactSteam Condensation with SimultaneousNoncondensable Gas Absorption, AIChE J., vol. 41,

No. 2 (1995), pp. 241 2507. Henriques de Brito M., von Stockar, A., Menendez

Bangerter, A., Bomio, P. and Laso, M., EffectiveMass-Transfer Area in a Pilot Plant Column Equippedwith Structured Packings with Ceramic Rings, Ind.Eng. Chem. Res., vol 33 (1994), pp. 647-656

8. Karapantsios T.D., Kostoglou M. & Karabelas A.J.,Direct Contact Condensation of Dilute Steam/AirMixtures on Wavy Falling Films, Chem. Eng. Comm.,vols. 141 142, (1996), pp. 261 285

9. Fair J.R, & Bravo J.L., Distillation columnscontaining structured packing, Chem Eng. Progr. , vol86(1), (1990), pp. 19-29

10. Schpigel L., Meier W., Performance characteristics ofvarious types of Mellapak packings (productivity,pressure differential, and deficiency), Chemical andPetroleum Engineering, vol 30, No. 3-4, (1994), pp.118 125


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Direct Contact Heat Transfer