THE PERFORMANCE OF TUBULAR AND DIRECT CONTACT JUICE HEATERS

P. G. Wright Sugar Research Institute, Mackay, Queensland, Australia

ABSTRACT

As a result of the considerations of the fundamental correlations for heat transfer and pressure drop as applied to the operation of juice heating in raw sugar mills, <<new>> expressions are presented for the tubular juice heaters typically used. The performance of direct vapour contact juice heaters is also discussed.

INTRODUCTION

The process of juice heating, usually carried out in two stages denoted primary' and secondary heating, can be operated in a variety of ways, and use a range of steam or vapour sources. The Australian raw sugar industry has for various reasons standardised on the use of the horizontal cylindrical multipass tubular type, using vapour direct from the exhaust main, or from the first or second evaporator stages. In other countries, the use of vertical cylindrical multipass types with longer tubes is common. More recently, the present author has advo- cated the use of direct contact methods of juice heating (Wright '; Wright and Hutchinson lo) and has reported on the performance of such devices. However, the traditional multitubular heaters will remain for a long time, and design infor- mation on these is still desirable.

TUBULAR JUICE HEATER PERFORMANCE

It must be concluded from a review of the present literature that the many equations given for prediction of heat transfer coefficient (HTC) give a wide scatter of predictions for typical heater arrangements, and are generally too sim- plistic to be accurate. However, since in the Australian cane sugar industry the design of juice heaters is very standardised, and the operating conditions do not vary widely, it is possible to formulate design expressions for heat transfer and pressure drop which are quite accurate. As the complexity of an expression is now not a great barrier to its use, it seems opportune to review the expressions available, and to look more *closely at the fundamental chemical engineering approach to the formulation of heat transfer and pressure drop equations to see if more accurate expressions can be derived. This has been done recently by the present author (Wright 14), and this paper summarises and extends the results.

Review of Design HTC Estimates

A short summary of the range of heat transfer expressions for the cane juice heating application has been given by Hugot. If these are restricted to practical operating heaters with a 72 to 144 h cleaning cycle, the expressions which could be used are as follows:

1. The Hugot Formula

U = 0,00554TS V(V + 0,04)

where the symbols and units are given in the symbol table.

2. The Australian Formula (Anon l)

U = 1,028 fl with the same nomenclature as above. This equation was formulated from data collected for the QSSCT group during the 1933 season.

3. The Crawford and Shann Formula (Crawford and Shann 3,

which Hugot recognised as having a more correct form that the preceeding expressions, but which gives values for ((clean tubes), only.

4. The Modified Hugot Expression (Hugot 4,

which retains a steam temperature variation factor and a modified velocity term.

When the formulae above are evaluated and compared for typical practical Australian heater conditions they give very wide differences. Of those given, the ((Australian formula. (Item ' above) is considered by the author to give a reasona- ble basis for sizing juice heater installations which use a 72 h cleaning cycle. Ho- wever, the equation does not predict changes of HTC with time ((on line>> due to scaling influences, and is greatly oversimplified.

~undahental HTC Considerations

The fundamental expressions relating to heat transfer in tubular heat exchan- gers are available in many standard texts on heat transfer and chemical enginee- ring unit operations. In view of the similarity of heater design and of material properties in the Australian sugar industry, it is possible to narrow the examina- tion to that for standard designs of heaters, with horizontal or vertical layout, with brass or stainless steel tubes, ant to consider the juice as equivalent in viscosity and thermal properties to a 20 brix sucrose solution. This simplifies the task very significantly.

The Overall Heat Transfer Coefficient of the Heater

The overall heat transfer coefficient U ( ~ W / ~ ~ . " C ) , of the heater determines the heat flux generated in the heater for a given temperature force, and is thus the most important performance factor for the unit. The overall coefficient is related

1030 -

to the various individual transfer coefficients encountered in the process by the ' expression:

which gives the reciprocal of U as an overall <<resistance,) which is the sum of the individual ccresistances~ due to the steam condensation outside the tubes, the juice film inside the tubes, the tube material, and the scale layer respectively. The overall expression above results from the integration of the heat transfer expres- sion for a small element of the heater tube, the heat transfer process of which is illustrated in Fig. 1 .

The evaluation of the four terms in the overall heat transfer expression has been given in detail by Wright. l4

For the juice film resistance the expression of Sieder and Tate is used with the juice properties taken as for a 20 Bx sucrose solution, and the solution viscosity, correlated with film temperature. The condensate film resistance for horizontal or vertical layout is taken from McAdams expressions in Kern, and the metal tube resistances for tubes of 1,62 mm for brass and 1,2 mm for stainless steel, which are typical thickness for the application. A special correlation for the scale film resistance is obtained from the data of Crawford and Shann. This includes a fouling multiplier factor and is quite sensitive to the average juice velocity (V, mls) in the tubes.

Overall Heat Transfer Expression The consideration of the different terms comprising the overall expression for

the heat transfer performance of tubular juice heaters, as carried out above, can be brought together to arrive at overall heat transfer expressions as follows:

where the values of X1 to X5 are given for the different applications as follows:

Application X Values

x I x, x3 x4 x5

Primary, horizontal, brass Primary, horizontal, s . steel Secondary, horizontal, brass Secondary, horizontal, s , steel Primary, vertical, brass Primary, vertical, s. steel Secondary, vertical, brass Secondary, vertical, s. steel

---

The value of the scaling multiplier factor X4 has been found to be around 0,0035 in several cases, but can vary from 0,002 in low fouling primary heating applications in some factory areas to 0,005 in some secondary heating applica- tions. This has to be ccmatchedn to the typical factory heater performance pattern for greater accuracy.

3 L 0' a, ?= $ - .' b = a, w .2 r m m L m 3 V) - I- m - - c . 2 m a, g 6

0 E m , ' , U c

0 . 5 . - z c n 2 0 0 Metal wall

Distance

Figure 1. Temperature gradients for heat transfer in juice heating.

Figure 2 shows HTC's plotted against operational time from the expression (d), for secondary heating with horizontal 0,044 m diameter stainless steel tubes and a scaling multiplier (X4) value of 0,0035. The values predicted by the <<Australian formula>> agree with the 72 h operation values for the middle juice velocities in the range. The sensitivity of the <<new>> equation to juice velocity is more than that of the <<Asustralian formula>>, and this agrees with the contention of Crawford and Shann that juice velocity is important both from the juice film coefficient and from the rate of scaling aspect. The <<new>> predictions for clean tubes correspond quite closely to the <<clean>> data of Crawford and Shann.

0 10 20 30 40 50 60 70 80 90 100 110 T~rne of operation (h)

Figure 2. Heat transfer coefficient prediction for secondary juice heaters.

Also shown in Fig. 2 is a plot of the HTC of a large secondary juice heater determined at Plane Creek factory for one week in November, 1980. The agree- ment to the 1,5 mls velocity prediction curve is excellent after the first period of reduced juice flow on start up has passed. (The necessary data was logged using a logic controlled data logging system based on Intel SBC 80120-4 microcomputer boards.)

~ ~ ~ l i c a t i o n s of HTC Expressions

Single Heater Expressions

Where a single heater, or group of heaters used together, is operated, the HTC can be calculated by the equations above as a function of time. An estimate of the HTC actually required to carry out the heating duty can then be made using the equation:

Q Ureq =A LMTD

Where U,,, is the HTC required for the duty, and LMTD is the <<log mean temperature difference>> driving force for heat transfer across the heater, and A is the heater total area. The heat load, Q, can be calculated from the juice flow rate (JF) and the temperature rise in the juice (Tout - Tin) by the formula

The "log mean temperature difference" can be calculated from the saturated steam temperature (T,) and the juice inlet and outlet temperatures by the formula:

LMTD = Tout - Tin - m

2,303 lg '" lin

Ts - Tout

Obviously, the single or grouped heater can only be run until its HTC drops to the level of the HYC required for the job, after which it must be supplemented by another heater, or relieved of the duty while cleaning is scheduled.

Multiple Heater Systems

Where a number of heaters are operated, with some available for mid-week cleaning, the prediction equations should be useful in evaluating possible schedu- les for operation. It is not necessarily best to have all available heaters <<on line>> early in the week, as it is seen that higher juice velocity tends to reduce the rate of scaling to assist the situation at the end of the operating period.

Juice Heater Pressure Drop Expressions

Design expressions for the pressure drop across tubular juice heaters are useful for an estimation of juice pump duty, and for the estimation of the maximum juice velocity which can be tolerated by a juice pump installation. The one most commonly used has been given by Hugos, and this relates the pressure head across a heater to the velocity of juice in the tubes, and the number and length of the juice passes through the heater as follows:

where Pt is the pressure head loss in metres of juice. The Hugot equation involves some simplifications and assumptions, and it was

considered worthwhile to look again at the fundamental dimensionless correlations for pressure drop in juice heaters, to see is improvements can be made.

Tube Pressure Loss

Using basic correlations for the pressure drop through tubes by Sieder and Tate given by Perry ' and making some simplifying assumptions for wall viscosity effects and fanning friction factor, Wright l4 has obtained a new expression for the pressure loss in the tubes. Adding this to the flow reversal loss at the end of each pass as estimated by Crane and Perry, ' the total pressure loss in metres head of juice is given by

Pt = O , O O ~ ~ ~ ( D V T ) - ~ ~ ~ V ~ ~ + 0,1018 n v 2 (m head juice)

This can be considered a "new" pressure drop expression, which is more exact than that of Hugot quoted above.

A comparison of the predictions for pressure head loss by the Hugot expression and the "new" expression abbve is given in Table I for a typical secondary juice heater with a range of velocity values. It is seen that the "new" expression giver slightly higher results in this case. The differences between the "new" expression and the Hugot expression are greater with long tube heaters, the Hugot values being too high.

Table 1. Comparison of Predicted Pressure Head Losses Throug a Juice Heater

Velocity of Prediction Prediction Juice in Tube Hugot (1972b) this Work

(m/s) (m) (m)

1 ,o 2,9 3,4 1,s 6,6 7,4 2,o 11,8 12,9 2,s 18,4 19,8 3 ,o 26,5 28,l 3,s 36,O 37,9 4,o 47,O 49,l

' Tube diameter 0,051 m; passes 20; length 3,O m; average juice temperature 85°C.

DIRECT VAPOUR CONTACT HEATER PERFORMANCE

General

A paper by Valdes 'l to the ISSCT conference in Brazil drew attention to the direct contact methods of juice heating using vapours from the evaporator set. The advantages and disadvantages of this method have subsequently been explored by the present author (Wright, Wright and Hutchinson lo), and the performance of a full factory application described by Young and Westmoreland. l2 Guidelines for the sizing of contactors have recently been presented by Wright. l3 A number of Australian factories have installed such devices for primary juice heating since 1980.

Contactor Performance

Contactor performance can be considered from two angles (a) the general equipment sizing to achieve a satisfactory "approach temperature" (here defined as the difference between the saturated vapour temperature of the steam source and that of the juice leaving the contactor), and (b) the process steam economies and evaporator loadings generated by the contact heating application.

The equipment sizing can be made to follow the general design expressions for a rain tray condenser as given by Wright, l3 with the modification that the water flow W is replaced by the anticipated juice flow to be treated. The vapour flow can then be estimated for the anticipated juice temperature rise and flow as for a normal heater. An "approach temperature" performance of less than 3°C should be possible if these guidelines are followed. In pilot plant trials described by Wright and Hutchinson both conventional and modified "doughnut-disc" rain

tray internals were used, and volumetric capacity loadings of 100 t/h juice per 1,O m3 of contactor body volume were handled with an ccapproach temperature,, less than 1°C. Measured factqry scale results have not so far been quite as good. No problems with juice frothing or blockage on the tray have been en- countered so far, this being assisted by efficient prescreening of the juice strea- m. It is, of course, important that the air entering with the juice flow be minimised by proper level control on the juice inlet tank.

Direct vapour contact results in the condensation of the heating vapour flow (VF) direct into the juice, and, as this must ultimately be removed in the evapora- tor station, the total evaporation requirements of the station increase. As discus- sed by Wright, this re uires an overall increase in multiple evaporator area of 9 approximately 33 VF m for a typical quintuple set, or of 27 VF m2 for a typical quadruple set where the units of V F are tonne per hour.

Direct contacting also means that, if vapour be taken from the same source stage of the evaporators as for conventional heating, the overall process steam economy decreases by VF/N where N is the number of evaporator stages. Howe- ver, because of the low "aproach temperature" performance possible whith a direct contactor, it is feasible to take vapour from a source significantly further down the set. As given by Wright, ' for every stage further down the set that the vapour for direct contacting vapour flow VF is selected, an approximately steam saving of VF/N is achieved. Thus, in typical Australian practice, for pri- mary juice heating shown in Fig. 3, direct contact vapour bleed from No. 4 stage of a q u i n t u p l e s e t wou ld give a n ove ra l l p rocess s t eam saving of

1 3 2 - ?VF + ,VF = , VF over the conventional use of No. 1 vapour with tubu- J J J

lar juice heaters for primary heating. The flow of final stage vapour to the evaporator condenser is reduced by a similar amount.

The shifting of heater bleed further down the multiple effect set does cause a change in the evaporative loading of the individual evaporator stages. As discus- sed by Wright ' the evaporative loadings of each evaporator stage "bypassed" by the change in heating source increase, and that of the remaining stages before and after the "bypassed" stages are reduced. It can be shown that if M is the number of stages "bypassed" in a multiple set, the reduction in evaporative loading for

' each remaining stage is given by VF(M-1)/N while the increase in the evaporative loading of the "bypassed" stages is given by VF(N+l-M)/N. The sum of the increases and decreases through the set gives an overall evaporation increase of VF for the entire set as mentioned before. In the example before of vapour bleed shift from the No. 1 to No. 4 stage of quintuple set, N = 5 and M = 3 so that the overall evaporation rates relative to the original case - 2VFl5, 3VFl5, 3VF/5, 3VFl5 and - VFl5 respectively for stages 1 to 5. This results in the percentage changes shown in Fig. 3.

1 Contartor Control

Several methods of oulet juice temperature control have benn used with direct contact juice heaters. The pilot experiments of Wright and Hutchinson lo used a temperature sensor in the outlet leg to position butterfly valve in the va our inlet ~ line. The factory installation described by Young and Westmoreland l4 controls temperature by throttling the gassing line leading to the vacuum pump. This

I gives a slower, but yet satisfactory, control with a relatively small control valve. Other factory installations use a controlled bypass of unheated juice from the inlet line direct to the outlet leg, with the temperature sensing of the juice after mixing

ii: .L I . .

is allowed to take place. As long as the source vapour is of good sufficiently steady pressure, then outlet temperature control has been found to be very easy to achieve.

DISCUSSION

The expressions presented above for tubular juice heater coefficients and pres- sure drops should be preferable to the simple expressions currently being used, as they take into account more of the variables associated with the operation, and follow the accepted chemical engineering approach more closely. Their higher complexity is not a problem, as the expressions can be readily solved with modern calculators and computing aids.

The performance of direct vapour contact juice heaters in related to their "approach temperature" and is dictated by the design of the contactor. This should follow similar guidelines to those arrived ar for water condenser design. The implications of direct vapour contact heating with respect to process steam economy and multiple effect evaporator capacity can be followed by use of the simple expressions given here. More exact calculations can be made using sophis- ticated computer simulations of multiple evaporator sets, but the approximations given should be satisfactory for most purposes.

SUMMARY OF TERMS

Symbol Description and Unit

A C D Hc H f

.HR JF Kt Ksc

LMTD L N n M Pt Q

Total heat transfer area of the heater. Specific heat of the juice, kJ/(kg."C) Internal diameter of the heater tubes, m Heat transfer coefficient through the condensate film, kW/(m2, "C) Heat transfer coefficient of the juice film, kW/(m2."C) Time of operation since cleaning, h Juice flow rate to the heater, kglh Thermal conductivity of the tube metal, kW/(m."C) Thermal conductivity of the scale layer, kW/(m."C) Logarithmic mean temperatpure difference for the heater, " C Length of heater tubes, m Number of effetc stages Number of passes through the heater Stage from which direct contact vapour bleed is taken Pressure head loss throught the tubes of the heater, m juice Heat flux for the heater, kW Steam side temperature, "C Temperature of the juice, "C Juice temperature leaving the heater, "C Juice temperature entering the heater, "C Thickness of tube wall, m Thickness of the scale layer on the tube, m Overall heat transfer coefficient (HTC) for the heater, kW/(m2."C) The overall heat transfer coefficient required for a heater duty, kW/(m2."C) Juice velocity in the tubes, m/s Vapour from rate to bleed heaters, t/h Values in the overall heat transfer expression

1038

REFERENCES

1. Anon (1936): International Sugar Journal, p. 438. 2. Anon (1969): Crane Valve Co. Technical Paper No. 410.A, p. 27. 3. Crawford, W. R. and Shann, D . S. (1956): Heat Transfer in Juice Heaters. Proc. Int. Soc., Sugar

Cane Technol. 9th Congress, India, p. 103. 4. Hugot, E . (1972): Handbook of Cane Sugar Engineering, Elsevier, 2nd edition, pp. 450-451. 5. Hugot, E. (1972b): Handbook of Cane Sugar Engineering, Elsevier, 2nd edition, p. 445, equa-

tion 31.17. 1 6 . Kern, D . Q. (1950): Process Heat Transfer. McGraw Hill. New York. 1st edition, p. 266,

equation 12.44a. 7. Perry, J . H.(1963): Chemical Engineering Handbood. McGraw Hill. New York. 5.33. Table 5.19. 8. Sieder and Tate, in Perry, J . H. (1974): Chemical Engineering Handbook.McGraw Hill. New

York, 5th edition, 10.14, equation 10.50. 9. Wright,P. G . (1979): Juice Heating by Direct Vapour Contact. Proc. Aust. Soc. Sugar Cane

Technol., 53rd Conference, p. 161. 10. Wrigth, P. G. and Hutchinson, R . T . (1980): A Plant Trial of a Direct Vapour Contact Juice

Heater. Proc. Aust. Soc. Sugar Cane Technol., 54th Conference, p. 17. 11. ValdCs, D. A. (1977): Heater Designs for the Cane Sugar Industry. Proc. Int. Soc. Sugar Cane

Technol., 16th Congress, Brazil, pp. 2401-2403. 12. Young, R . C. and Westmoreland, A . . (1981): Primary Juice Heating by Direct Vapour Contact.

Proc. Aust. Soc. Sugar Cane Technol., 55th Conference, p. 17. 13. Wright, P. G. (1982): Rain Tray Condenser Desisgn Notes. Proc. Aust. Soc. Sugar Cane Tech-

nol., 56th Conference, p . 273. 14. Wrigth, P. G. (1981): Performance Equations for Horizontal Tubular Juice Heaters. Proc. Aust.

Soc. Sugar Cane Technol., 55th Conference, p.135.

FONCTIONNEMENT DES RECHAUFFEURS DE JUS TUBULAIRES ET PAT CONTACT DIRECT

P. G. Wright Institut de recherches sucrikres, Mackay, Queensland, Australie

RESUME

Se basant sur les correlations fondamentales de transmission de'chaleur et de chute de vapeur appliquees B l'opera'tion de rechauffage du jus en sucrerie, l'auteur presente des expressions "nouvelles" pour les rechauf- feurs de jus tubulaires utilisks normalement. I1 se penche aussi sur la question du fonctionnement des rechauffeurs de jus par contact direct avec la vapeur.

RENDIMIENTO DE LOS CALENTADORES DE JUG0 TUBULARES Y POR CONTACT0 DIRECT0

P. G. Wright Instituto de Investigaciones Azucareras, Mackay, Queensland, Australia

RESUMEN

Como resultado de la consideraci6n de las correlaciones fundamentales sobre termotransferencia y caida de presi6n aplicadas a la operaci6n de calentamiento del jugo en las fibricas productoras de azucar crudo se exponen "nuevas" expresiones para 10s calentadores de jugo tubulares tipicamente utilizados. Se discute, ademis, la operaci6n de 10s calentadores de jugo por contact0 direct0 del vapor.