Pitfalls in ammonia absorption refrigeration

Pitfalls in ammonia absorption refrigeration M. J. P. Bogart

Key words: ammonia, absorption refrigeration, refrigeration

Les problCmes des syst mes frigorifiques & absorption d'ammoniac Le syst#me frigor/fique ~ absorption ammoniac-eau a #td invent# il y a 131 ans, mais compte tenu de sa

Aqua-ammonia absorption refrigeration (AAR) was invented 131 years ago, but this versatile refrigeration process is increasing in importance today since it consumes mainly

grande souplesse d'utilisation, i/ prend de /'importance actuellement du fait qu'il consomme surtout de I'dnergie thermique ~ bas niveau. On analyse ce syst#me et on examine ses prob/#mes de conception et de fonctionnement au niveau de I'dvaporation, de/'absorption et de la distillation.

low level heat energy. The AAR technique is analyzed and its design and operating problems are discussed with reference to evaporation, absorption and distillation.

Nomenclature

Subscripts

K L N P R t T V X Y /~

c critical state E refrigerant evaporator r reduced state R refrigerant t theoretical tray W water 0, 1, 2 . . . n - 1 , n tray number

equilibrium vaporization ratio, Y/X liquid flow, kg h -1 number of trays absolute pressure, kPa refrigerant flow, kg h -1 temperature, °C absolute temperature, K vapour flow, kg h -1 liquid-phase composition, ppm (wt) H20 vapour-phase composition, ppm (wt) H20 b l o w - d o w n rate/ w t liquid entrainment rate/ wt dry vapour

Aqua-ammonia absorption refrigeration (AAR) was invented 131 years ago as a batch process. Its extensive modernization has not yet reached a peak, probably due to the lack of attention on the part of chemical engineers. This versatile refrigeration process is becoming increasingly attractive as energy costs continue to s o a r - its major advantage is that it consumes mainly low-level heat energy. The author is from the Advanced Technology Division, Fluor Engineers and Constructors, Inc., Irvine, California, USA. This paper was presented at the 2nd World Congress of Chemical Engineering, Montreal, Canada, October 6, 1 981.

Volume 5 Num6ro 4 Juillet 1982

Superscript o

pure ammonia

Use in a m m o n i a m a n u f a c t u r e The economic gain inherent in the AAR technique has been incorporated into the Fluor Ammonia Process. ~ Here carbon dioxide is removed from the raw synthesis gas by physical solvents rather than aqueous amine or carbonate solutions; the heat energy thus relinquished drives an AAR unit which provides the total refrigeration requirements of the ammonia plant. The saving in natural gas consumption over the conventional mechanical (compression) refrigeration route is reported 2 to be upward of 6 0 0 0 0 0 Btu (LHV) per short ton of

01 4 0 - 7 0 0 7 / 8 2 / 0 4 0 2 0 3 - 0 6 8 3 . 0 0 © 1982 Butterworth & Co (Publishers) Ltd and IIR

203

ammonia. In a 1200 ST/D ammonia plant, with natural gas purchased at $3.00 per million Btu, the annual reduction in operating costs is $720000. The savings in capital costs are about equal to this sum; the total gain over conventional technology is then about $3.60 per short ton. This novel ammonia process has been commercially proven in two 1200 ST/D plants which have been in operation since 1977. These plants have met or exceeded design performance. Other benefits include: One, moving parts are limited to standard pumps and control valves. Hence, maintenance costs are reduced, and down-time revenue loss is minimized. Two, plant start-up time can be shortened by precooling the low-temperature equipment and circuits using stand-by low-pressure steam. Three, the AAR unit functions as an independent heat sink. It readily accepts a varying heat input and load (eg, summer/winter, daytime/night-time changes) without upsetting the plant operation. It eliminates the start-up problem of cutting in the refrigeration compressor turbine without unbalancing the complex steam system. Four, partial-load operations are easily accommodated since the heating requirement remains proportional to the load. The turn-down ratio is limited only by the design of the fractionator trays; it is not constrained by the ~50% surge point of centrifugal compressors. Five, transient liquid slugs or carryover from the refrigerant evaporators (chillers) will not physically damage downstream equipment.

105 ° k.

103 -5(

-40

-30

-20 -10 Temperature, °C

\

O

10

20

Fig. 2 Phase-equilibrium relationships foraqua-ammonia evaporators Fig. 2 Rapport entre/es teneurs en eau ~ /'#qui/ibre entre/es phases/iquides et vapeur dans un #vaporateur ammoniac-eau

areas that might seriously impair the operation and performance of their designs. It is the purpose of this paper to point out such major pitfalls and to present means for avoiding them.

A A R functions The AAR technique will be analysed and its design and operating problems discussed according to the functions of its basic elements:

Basic A A R unit

Evaporation

Inspection of the process flow diagram of a basic AAR unit, Fig. 1, shows that it is a simple chemical plant working with a classical binary system, whose thermodynamic properties are well-known. Its apparent lack of sophistication could well beguile the novice and expert alike into overlooking problem

The desired refrigeration effect is conventionally produced by boiling off refrigerant in the evaporator. Once the process conditions are properly set and the form chosen (kettle-type, vertical thermosyphon, etc.) completing its design becomes a conventional task for the heat-exchanger specialist.

Overhead , ~ c ° n d e n r ~

A.,

Mixer

An initial appraisal of the distillation task suggests that it should be easy to produce the refrigerant ammonia as overhead distillate with a water content in the Iow-ppm range. A serious error may be made by ignoring this contaminant water and using the properties of anhydrous ammonia for the process design.

~

*oo"7g

Strong - aqua Reboiler

uD

Preheater

:~

Fig. 1 Basic flow diagram, aqua-ammonia absorption refrigeration unit Fig. 1 Schema de principe. Insta//ation frigorifique ~ absorption d'ammoniac-eau

204

The pertinent phase equilibrium data are summarized in Fig. 2. Taking, as an example, an evaporator with a -30°C refrigerant boiling temperature, it is seen that the equilibrium water content of the liquid phase may approach 20000 times that of the vapour phase. Consider the simplest evaporator as the heat exchanger of Fig. 3 and ignore, for the moment, entrainment and blow-down. The effluent vapour is then identical in flow and composition to the refrigerant stream supplied to the evaporator. If this ammonia supply contains only 10 ppm of water, the boiling liquid will approach an equilibrium water content of 10 x 17 000=170000 ppm.

International Journal of Refrigeration

The process conditions for an evaporator with such a diluted refrigerant liquid pool will be far from those for anhydrous ammonia. Normally, the process demand sets the refrigerant boiling temperature as inviolate. In this example, the calculated evaporation pressure will then drop from 119 kPa for pure ammonia 7 to ~ 9 7 kPa for a boiling liquid of only 83 wt% NH 3. As may be deduced from Fig. 1, this will decrease the efficiency of the AAR unit by lowering the absorption pressure, delivering a more dilute feed to the fractionator. We have also lost the possibility of avoiding vacuum operation. This situation is aggravated in a multi-level cascade wherein the refrigerant flows in series from the highlevel evaporators to the low-temperature ones. Very little water will be rejected in the vapour from the higher level units, nearly all of the water in the refrigerant supply will be concentrated in the feed to the end-of-the-line (lowest temperature) evaporator. This water content can thus be several times that of the supply and lead to an unacceptable lowering of evaporator pressure. The key to this problem lies in purging by blowdown the water which accumulates in the liquid of the terminal evaporator. This is partly and involuntarily done by the liquid droplets entrained from the vigorously boiling liquid into the effluent vapour. However, entrainment can vary unpredictably and may not give sufficient relief. The resulting evaporator liquid concentration depends on the sum of the entrainment and blowdown rates ( # + ~) it is given by: 3 XE----XR(1 + ~ +

s)/(Kw+l~+

E)

(1)

The evaporator pressure deviation from that for pure ammonia may be determined from the phase equilibrium data. For this purpose it is.usually sufficient to calculate it from the approximate relationship

Dry

= VE, YE

vol:x3ur

EntPmnment EVE,XE

"

PE I

Refrigerant supply Hot medium ~ [ ~ i ? i i i i ~ i i ~ : ~ J J

-

R XR

B l o w - d o w n ~ VE,XE

Fig. 3 Schematic diagram showing refrigerant evaporator (chiller)

Fig. 3

[email protected] frigorifique (refroidlsseur)

'2°1 \

ur

oot a°l-

\

oo4oL : O -2C O Fig. 4

\

\ T4-

v

i %L oi' ~ #

Vo ~Setureted

275 kPa I I I I I I I 20 40 60 80 Composition, weight percent ammonia

liquid I 1OO

Isobaric t-x-y diagram for ammonia-water at 275 kPa

Fig. 4 Diagramme @pression constante t-x-y pour/'ammoniaceau @275 kPa

Absorption

Fig. 4 is the isobaric phase-equilibrium diagram for the system ammonia-water at a representative evaporator pressure, 275 kPa: superimposed on it is the proper operation of an absorber-condenser. The feeds to the heat exchanger in this example are ,-,100% NH 3 vapour (R) at 20°C and 15 wt% NH 3 weak-aqua (point A) at 40°C. These streams are premixed upstream (Fig. 1), the heat of solution will raise this blend (bulk composition 42 wt% NH3) to 66"C. This entering liquid-vapour mixture is represented by point M in Fig. 4. If the two phases are kept in intimate contact, they will cool along line MS until they reach the liquidus at point S, at 35°C. Further heat exchange will simply subcool the condensate.

The vaporized refrigerant is absorbed in water or weak-aqua. If care has been taken to exclude noncondensible gases, we have no need for the countercurrent stagewise absorber, which separates a solute gas from insoluble ones. The task is to dissolve ammonia vapour in water and to remove the considerable heat of solution generated.

The exchanger best suited for this service is the vertical wetted-wall, falling-film type. The liquid and vapour phases are here maintained in good contact with each other, promoting simultaneous mass- and heat-transfer. Air-cooled units also perform well, thanks to a number of return bends which repeatedly remix the two phases.

APE----100(PE° - PE)IPE ° =XE( Tr2/3)/7500

(2)

where T,= TE/Tc, the reduced state temperature of the evaporator liquid. Most evaporator water-accumulation problems are satisfactorily resolved with a blow-down rate of 1 to 5% of the total refrigerant supply quantity. In evaporator cascades the value of # in (1) will be higher by the ratio of the total refrigerant flow to the cascade to that fed to the terminal unit.

Volume 5 Number 4 July 1982

205

50(

/ Saturated vapour

40

[_ O

E

30

Q.

20

the minimum. This is due to of the two components and span involved: 44 ° to 205°C fractionation pressure, 1725

1725 kPa Enthalpy of pure liquids =O at O°C

The large AH v variation helps explain w h y the McCabe-Thiele method erroneously gives a minimum reflux ratio that can be as low as one-half to one-quarter of the true value. 4 It also fails to warn that the true minimum reflux ratio value may increase by 50% or more as the overhead distillate composition is varied over the very narrow range of 99.5% to ~ 1 0 0 % NH 3. The sharp rate of change of AHv at the high-ammonia end appears to be the cause of these deficiencies.

Heat of vaporization

hi

10 p

ol 0

~

j S a t u r a t e d liquid

I 20

[

the highly polar nature to the wide temperature at a representative kPa.

I - -

40 60 Weight percent ammonia

I 80

I 100

Fig. 5 Isobaric enthalpy-concentration diagram for ammoniawater at 1725 kPa

Fig. 5 Diagramme ~ pression constante concentration-enthalpie pour I'ammoniac-eau ~ 1725 kPa

At the opposite end of the scale is the horizontal shell-and-tube unit with condensation taking place in the shell side. Such a design unfortunately tends to act as an efficient vapour-liquid separator, preventing the essential maintenance of contact between the phases. At the extreme, the vapour stream will cool along dew-point curve ViV o and the liquid along bubble-point c u r v e L i S (more likely, along the subcooled curve L,Lo).

The correct design data are obtainable from another rapid, convenient, graphical method which uses an enthalpy-concentration (Ponchon) diagram (such as Fig. 5) to determine reflux ratios and to correct for operating line curvature. The method is rigorous in that. it performs simultaneous mass- and heatbalances on the fractionator. Fig. 6 is an exploration by the Ponchon method of the two prime variables which set the minimum reflux ratio for the fractionator - the feed inlet temperature (regardless of feed composition, provided that it enters between its bubble- and dew-points), and the distillate composition. Note the increasing difficulty (added reflux and trays) of producing high-purity distillate refrigerant at the lower feed inlet temperatures.

Entrainment. Another adverse phenomenon is the effect of fractionator entrainment on the water 2.0 160°C Feed inlet

Distillation

The 'strong-aqua' absorbate from the absorption step is an ammonia-rich aqueous solution. Distillation is used to separate it into its component parts, which are recycled as shown in Fig. 1. The fractionator overhead distillate is the high-purity ammonia returned as refrigerant to the evaporator and at the bottom is the weak-aqua. Optimally, this weak-aqua does not need to be stripped relatively free of ammonia. A modest ammonia content effectively drops the bottom temperature in the fractionator reboiler and allows the use of a lowerlevel heat source to "drive' the AAR unit.

Ref/ux ratio. The McCabe-Thiele x-y diagram might be expected to provide the basic process design data for the fractionator (reflux ratio vs theoretical plates) with minimum effort and risk of error. Unfortunately, the simplicity of this classic method hinges on the equality of the molar latent heats of vaporization for the two components over the composition range involved to give straight operating lines for the rectifying and stripping sections. Fig. 5 shows that the ammonia-water system departs widely from this criterion. AH~ varies irregularly with composition; the peak value is twice

206

1.0 ~ _ ~

~

'

140°C

~

120Oc

O

100o C

_ 6 ° ° cC O.Ol 0

/

~

i'x

\.

,

,

L

,

x,\

2 4 6 8 Distillate w a t e r content, thousand ppm's (weight)

10

Fig. 6 Effect of overhead composition and feed preheat on minimum reflux ratio

F/g. 6 /nfluence de la composition en t#te et du pr#chauffage /'entr#e sur le taux minimal de reflux

Revue Internationale du Froid

|

Lo

R

ID

X Tr'ay 1

L1 Xl

System pressure

V2

r2

Fig. 7

A final caveat concerns the operating pressures in the AAR unit. The fractionator is at the higher pressure dictated by the temperature of the coolant available for condensing over-head vapour. The lowest pressure will be that of the evaporator generating the coldest refrigeration. These pressure extremes are not under the control of the AAR designer.

Schematic diagram s h o w i n g fractionator top section

Fig. 7 Schdma. Secdon sup#rieure de I'apparei/ de fracdonnement

content of the overhead distillate product. At typical fractionator top conditions, the tray liquid water concentration is about 400 times that of the vapour in equilibrium with it, suggesting an easy fractionation. This advantage, however, may be seriously "watered-down' by entrainment. Consider, as a possibly extreme purity example, producing a distillate liquid analysing 1 ppm H20. In the arrangement of Fig. 7, the classical theoretical tray concept (with Y1 =XR = 1.0) states that X 1= 1.0 x 4 0 0 = 4 0 0 ppm I-t20. Imagine the effect on the bulk composition o f stream V 1 of the carry-over of even a small amount of droplets or froth from this highly water-enriched tray liquid. Fractionator entrainment depends on system properties, process conditions, and the hydraulic and mechanical tray design. If an average entrainment rate can be set, the effect on the tray composition profile may be estimated. On tray n, the compositions of the dry (entrainment-free) vapour, Yn, can be related to that of the liquid entering from the upper tray, Xo_ 1, by the approximation 5

(L/V+ ~/2)(Xn_l) -'}-(1 -L/V)(XR) Yn1+(~/2)/(Kw) n

Vacuum operations will occur in the AAR unit whenever the lowest refrigeration temperature is

w o

31 7

,/'.,.°2" ,

,,07"

(3) hi

(4)

Equations (3) and (4) may easily be used in an iterative procedure to calculate the composition profile in the upper rectifying section of a fractionator where, say, Y < 1 0 0 0 0 ppm H20. Here the task is simplified as variations of L/V and s may be ignored. Note, however, that the value of (Kw) n may vary more than ten-fold over this zone.

Volume 5 Num6ro 4 Juillet 1982

The literature suggests the viability of this technique to temperatures as low as -60°C. The vapour pressure of ammonia at this temperature is 6.9 kPa - the minimum system pressure is further lowered by the evaporator temperature approach and by the correction for water in the refrigerant supply. In addition to the poor refrigeration efficiency of the AAR unit at this very low level, 6 the capital invested will be further increased by the large-size equipment required to handle the high volumetric flows.

"

Under the equilibrium tray concept, the composition of the liquid leaving tray n is Xn=Yn/(Kw) n

An illustration of the magnitude of the effect of entrainment thus calculated is presented in Fig. 8. This graph shows the number of theoretical trays that must be added, (ANt), to the fractionator to reduce the tray vapour composition from an arbitrary intermediate Yn=10000 ppm H20 (1 wt%) to the specified 1 ppm H20 distillate product purity. The normal range of reflux ratio is explored, with the entrainment rate (ratio of liquid to dry vapour) varied up to the region of badly-designed or overloaded columns. As may be deduced, entrainment may be a prime cause for the failure of a fractionator to produce specification distillate purity.

0 2

/~,'.Ir i ~ 1 4 6 8 10 AN t = Additional theoretical trays

1 12

14

Fig. 8 Effect o f entrainment on tray requirement in ammoniawater fractionators

Fig. 8 Influence de/'entrafnement sur le nombre de plateaux clans/es appareils de fractionnement d'ammoniac-eau

207

below the normal boiling point of ammonia, - 3 3 ° C . Unfortunately inleakage of atmospheric air is almost impossible to prevent in such cases. The quantity ingested will not normally be enough to enter the narrow flammability-limit zone for ammonia-in-air. However, carbon dioxide is also introduced to react with aqueous ammonia, with the possible production of corrosive ammonium carbamate. This threatens the AAR advantage of being able to avoid materials of construction more exotic or expensive than carbon steel. Monitoring the day-to-day operation of AAR units with sections at below atmospheric pressure is certainly the very least that can be done to contain this problem.

208

Much of the material presented here is abstracted from a recent book on absorption refrigeration, 3 with the kind permission of Gulf Publishing Company.

References 1 Bogart, i . Jo P. Processfor ammonia manufacture, US Patent 3743699 (July 3, 1973) 2 Sogart, M. J. P. Save energy in ammonia plants Hydrocarbon Processing (April, 1978) 145 3 Bogart, M. J. P. Ammonia absorption refrigeration m industrial processes, Gulf Publishing Co., Houston (1981) 103 4 Ibid, 197 5 /bid, 217 6 /bid, 366 7 Haar, L., Gallagher, J, S. Thermodynamic properties of ammonia, J Phys Chem Ref Data 7 3 (1978) 635-792

International Journal of Refrigeration