# Experimental analysis of oxygen stripping from feed water in a two stage jet cum tray type deaerator.

IntroductionBoiler feed water may contain significant amounts of dissolved oxygen, which causes pitting and iron deposition. Deaerators are used to remove these dissolved and corrosive gases from boiler makeup feed water. The makeup water and the condensate are heated in the deaerator for removal of oxygen before entry into the boiler as even small amounts of dissolved gas can cause significant corrosion. The high temperature of boiler feed water will further increase the corrosivity due to dissolved oxygen, if left untreated. In Jet cum tray type deaerator, the incoming water enclosed in a chamber is passed through a pipe which is located at the top of the deaerator and gets collected at a distance of 0.55 m from the pipe, referred to as the first stage. The length of the second stage can be adjusted to 0.65/1.25/1.55m by using spool pieces between the trays. Both the first and second stages are engulfed with steam at a pressure. Experiments are conducted at different flow rates of deaerator water, lengths, deaerator pressures, inlet water temperatures, and oxygen concentration in inlet water.

Hasson et al. [1964a,b] theoretically analyzed the jets of various configurations such as fan spray sheet, uniformly thick sheet and cylindrical jet subjected to zero surface resistance. For Graetz number less than 30, the local Nusselt number is constant for all the three jets. In this region, the Nusselt number of cylindrical jet is observed to be lower compared to the other two. For Graetz numbers greater than 100, the local Nusselt number of fan spray sheet is observed to be higher compared to the other two types. The experimental investigation revealed the heat transfer coefficients to vary between 11630 and 23260 W/([m.sup.2]K) and remained unaltered when steam pressure is reduced from higher atmospheric to sub atmospheric pressure. The heat transfer coefficient decreases with increasing air content in steam. Nosoko et al. [3] conducted experiments on oxygen absorption using a single column horizontal tube bank of 16 mm diameter and 284 mm wetted length. They observed Sherwood number to increase with increase in tube spacing from 2 to 5 mm and then levels off at 10 mm or higher. They concluded that the volume of horizontal tube absorber to be 1/2.2 to 1/1.18 times lower compared to vertical orientation for the same heat duty.

The estimation of condensation heat transfer coefficient from jets by inducting non-condensable gas into vapor region has been analyzed by many. Heat and mass transfer studies with non-condensable gas such as oxygen getting stripped from boiler feed water under jet flow is quite limited. Hence, it is proposed to conduct experiments to estimate the heat and mass transfer coefficients at different mass flows rates and study its impact on oxygen stripping from the boiler feed water.

Fabrication of the Experimental setup

The experimental setup consists of 0.15m dia column and 1.2m long with flexibility to enhance the height to a maximum of 2.1 m using spacers. Spacers of 200 and 300 mm length are available which can be used individually or in combination to vary the length of the deaerator column. A 200 liters feed water storage tank, a steam jacket on the deaerator water inlet pipe for regulating its temperature and a pump for circulating water are other accessories. In the steam circuit a pressure regulator and steam trap are connected to a buffer tank for removal of water droplets after steam expansion in the pressure regulator. A water bath of 25 liters capacity with a copper coil to regulate the temperature of sample water to 30-40[degrees]C connected to Dissolved Oxygen (DO) meter, flow components such as valves, flow meters, pressure gauges and thermocouples are provided. The process and instrumentation diagram of the experimental setup is shown in Fig. 1.

[FIGURE 1 OMITTED]

Jet Flows

The estimation of heat transfer coefficients due to condensation of steam on water jets is of great significance in the design of equipment such as direct contact condensers, desuperheaters and for water desalination with "vapors reheat flash process". The liquid emanating as a jet flows for a short distance and later breaks-up to form droplets. High heat transfer coefficients can be attained with these cylindrical jets. The analysis of the data is based on the following assumptions.

Assumptions

(a) The liquid jet moves in a medium of constant temperature.

(b) Jet velocity is assumed constant, neglecting the interfacial drag and reaction on the jet due to vapor initially at rest condensing on a moving liquid.

(c) The variation of properties such as [C.sub.P], [rho], k of water with temperature is negligible

(d) The temperature of water in the second stage is assumed to increase linearly.

Estimation of Heat Transfer Coefficients--Stage 1

The heat conduction equation for cylindrical jets can be written as

[V.sub.1] [delta][T.sub.J]/[delta]z = [alpha][[[delta].sup.2][T.sub.J]/[delta][r.sup.2] + 1/r * [delta][T.sub.J]/[delta]r] [1]

where [V.sub.1], the jet velocity is estimated based on volumetric flow rate of water and the diameter of the nozzle. The variation of non-dimensional temperature with Graetz number for condensing vapors on laminar liquid jets for the condition of zero resistance at the interface has been presented by Hasson et al. [1]

[theta] = 0.6915 Exp (-23.136/Gz) for [Gz.sup.[less than or equal to]] 10 [2a]

and [theta] = 1 -8 [square root of ([pi] Gz)] for Gz > 10 [2b]

where Gz = [V.sub.1] [D.sup.2.sub.1]/[alpha]z

From the definition of dimensionless temperature, the temperature of water at the end of first stage can be determined from the relation

[T.sub.J] = [T.sub.S] - ([T.sub.S] - [T.sub.i]) [theta] [3]

The average Graetz number [Gz.sub.AV] for the first stage can be evaluated by integrating

[MATHEMATICAL EXPRESSION NOT REPRODUCIBLE IN ASCII.] [4]

The average Nusselt number and heat transferred in the first stage can be estimated

[Nu.sub.AV] = 5.784 for [Gz.sub.AV] [less than or equal to] 10 [5a]

[Nu.sub.AV] = [square root of (([Gz.sub.AV]/[pi]/1-8/[square root of ([pi] [Gz.sub.AV]))] for [Gz.sub.AV] > 10 [5b]

[Q.sub.1] = [h.sub.AV] [A.sub.1]([T.sub.S] - [T.sub.i]) [6]

where [h.sub.AV] = [Nu.sub.AV] k/[D.sub.1] [7]

The energy balance equation for the deaerator comprising of the first and second stages can be estimated from the relation

[Q.sub.E] = m [C.sub.PL] ([T.sub.o] - [T.sub.i]) = [Q.sub.1] + [Q.sub.2] [8]

Estimation of Heat Transfer Coefficients--Stage 2

A mathematical treatment of condensation on laminar and turbulent liquid jets emanating from small diameter jets has been presented by Mochalova et al. [4]. They compared their analysis with the experimental data of Mills et al. [5] and presented an explicit solution for the estimation of heat transfer coefficient in terms of Reynolds, Prandtl, and Weber numbers in addition to other geometric parameters governing the flow. For laminar flow the heat transfer coefficient can be estimated when [Re.sub.H] [less than or equal to] 1500 given by

St = ([f.sub.1][f.sub.2][f.sub.3] -0.004 Ja)[([L.sub.2]/[D.sub.2]).sup.-0.8] and [h.sub.2T] = St [[rho].sub.2][V.sub.2][C.sub.PL2] [9]

where

[f.sub.1] = 1.25 x [10.sup.-2] -7.5 x [10.sup.-6] [Re.sub.H]; [f.sub.2] = 1.05 - Pr (8 x [10.sup.-3] -3 x [10.sup.-4] [L.sub.2]/[D.sub.2])

and

[f.sub.3] = 1.05 - [C.sup.3/4] [Re.sup.-2.sub.H] We + [L.sub.2]/[D.sub.2] (0.03-0.6 [C.sup.3/4] [Re.sup.-2.sub.H] We) for We < 2.5;

[f.sub.3] = 1.05 - [C.sup.3/4] [Re.sup.-2.sub.H] We (0.4 + 0.01 [L.sub.2]/[D.sub.2]) for We [greater than or equal to] 2.5;

The heat transfer coefficient T h2 estimated with Eq. (9) can be validated for small diameter liquid jets, if the surface area of the second stage can be determined. The surface area for heat transfer depends on the breakup length of the liquid jet. The length of the jet emerging from 1.8 mm diameter holes of the tray can be estimated using the empirical correlation of Celata et al [6] given by

[L.sub.hb]/[D.sub.2] = 29 + 1.9 [square root of (We)] [10]

The volume of water in the jet upto break-up length can be estimated from the geometry

[v.sub.j] = [pi][D.sup.2.sub.2]/4 [L.sub.hb] [11]

The spherical droplet formation takes place on the breakup of the liquid jet. The diameter and volume of the droplet formed on breakup can be estimated from the relations given by Hinze [7]

d = 1.89 [D.sub.2] ; [v.sub.d] = [pi][d.sup.3]/6 [12]

The number of droplets N formed from each jet in the second stage can be estimated from the relation

N = [v.sub.J]/[v.sub.d] [13]

The surface area of the each tray, [A.sub.t] for heat transfer is the sum of the surface areas estimated upto breakup length and the surface area of droplets formed from 154 jets given by

[A.sub.t] = 154x[pi] ([D.sub.2] [L.sub.hb] + [Nd.sup.2]) [14]

Similar procedure is adopted for the remaining two trays. As the jet height from the fourth tray is 0.05m which is less than the breakup length, the surface area exposed to steam is taken for consideration. Hence, the total surface area of the second stage considering all the four trays for the estimation of heat transfer area is given by

[A.sub.2] = 3 [A.sub.t] + 154 x 0.05 [pi] [D.sub.2] [15]

The heat transfer coefficient of the second stage can then be estimated from Newton's law of cooling given by

[h.sub.2E] = [Q.sub.2]/[A.sub.2]([T.sub.S] - [T.sub.J]) [16]

The heat transferred in the second stage of the deaerator can be estimated from the energy balance Eq. (8) and written as

[Q.sub.2] = [Q.sub.E] - [Q.sub.1] [17]

It can also be estimated from the heat transfer coefficients evaluated for the first and second stages given by

[Q.sub.T] = [h.sub.1] [A.sub.1] ([T.sub.S]-[T.sub.i]) + [h.sub.2T][A.sub.2]([T.sub.S] - [T.sub.J]) [18]

Regression equation has been developed with 85 data points useful for estimating the overall heat transfer coefficient given by

[MATHEMATICAL EXPRESSION NOT REPRODUCIBLE IN ASCII.] [19]

valid in the range 0.033 < m 0.125 kg/s, 120 < L/[D.sub.1] < 210 and 1.1 < [T.sub.S]/[T.sub.i] < 1.4 with an average deviation (AD) of 3% and standard deviation (SD) of 4%.

Estimation of Mass Transfer Coefficients of both stages

The following assumptions are made in the analysis to evaluate the quantity of dissolved oxygen removed

(1) Jet is cylindrical in configuration

(2) Diffusion occurs under non-isothermal conditions i.e. the major resistance for diffusion is within the liquid

(3) The resistance for diffusion of oxygen from the interface to steam environment is negligible.

The component continuity equation can be written as

[pi]/4 [D.sup.2] L [[rho].sub.L] dX(t)/dt = -[k.sub.L,J] [C.sub.A] [pi] D L [20]

Further in the evaluation of X(t) in Eq. (20), the mass transfer coefficient, [k.sub.L,J] should be known apriory. The empirical correlations of Mayinger [8] valid in the laminar-wavy, transition and turbulent flow regime for falling films are given by

[Sh.sub.J] = 2.24x[10.sup.-2] [Re.sup.0.8.sub.J][Sc.sup.0.5.sub.J] [21a]

valid in the range 12 [less than or equal to] [Re.sub.J] [less than or equal to] 70 and [Sc.sub.J] [greater than or equal to] 2.32 x [10.sup.4]/[Re.sup.1.6.sub.J]

[Sh.sub.J] = 8.0x[10.sup.-2] [Re.sup.0.5.sub.J][Sc.sup.0.5.sub.J] [21b]

valid in the range 70 [less than or equal to] [Re.sub.J] [less than or equal to] 400 and [Sc.sub.J] [greater than or equal to] 1.82x[10.sup.3]/[Re.sub.J] [21c]

for [Re.sub.J] [greater than or equal to] 400 and [Sc.sub.J] [greater than or equal to] 1.47x[10.sup.7]/[Re.sup.2.5.sub.J] [21c]

Eq. (20) is rearranged and solved in conjunction with Eq. (21) for the initial condition t = 0, X = [X.sub.i] to obtain

[MATHEMATICAL EXPRESSION NOT REPRODUCIBLE IN ASCII.] [22]

Oxygen diffusion from droplets

The mass transfer coefficient of a spherical droplet can be derived from the component continuity equation given by

[pi]/6 [d.sup.3][[rho].sub.L] dX(t)/dt = -[k.sub.L,d] [C.sub.A] [pi][d.sup.2] [23]

Further in the evaluation of X(t) of Eq. (23) the mass transfer coefficient [k.sub.L,d] should be known apriory. For flow past a single sphere, the mass transfer coefficient under forced and free convective conditions is given by the well established dimensionless equations of Steinberger and Treybal [9]

[Sh.sub.d] = [Sh.sub.0] + 0.347 [([Re.sub.d] [Sc.sub.d.sup.0.5]).sup.0.62] [24]

valid in the range 0.6 < [Sc.sub.d] < 4000 and 1.8 < [Re.sub.d] < 6 x [10.sup.5]

The initial Sherwood number [Sh.sub.0] can be evaluated from the relation

[Sh.sub.0] = 2 + 0.569 [([Gr.sub.d][Sc.sub.d]).sup.0.25] for [Gr.sub.d][Sc.sub.d] < [10.sup.8] [25a]

[Sh.sub.0] = 2 + 0.0254 [([Gr.sub.d][Sc.sub.d]).sup.0.333] [Sc.sup.0.244.sub.d] for [Gr.sub.d][Sc.sub.d]> [10.sup.8] [25b]

Eq. (23) can be solved in conjunction with Eqs. (24) and (25) for the initial condition t = 0, X = [X.sub.i] to obtain

[MATHEMATICAL EXPRESSION NOT REPRODUCIBLE IN ASCII.] [26]

where [D.sub.L] = (7.481x[10.sup.-16] x [T.sub.AV])/([[mu].sub.L][V.sub.A.sup.0.6]).

The Eqs. (22) and (26) are solved for different mass flow rates, deaerator lengths and inlet oxygen concentrations in the experimental range and the results presented.

A regression equation has been developed for the estimation of Sherwood number in the experimental range 0.033 < m < 0.125 kg/s, 120 < L/[D.sub.1] < 210 and 1.1 < [T.sub.S]/[T.sub.i] < 1.4

[MATHEMATICAL EXPRESSION NOT REPRODUCIBLE IN ASCII.] [27]

obtained with AD of 2.54% and SD of 3.84%.

Results and discussion

The temperature and concentration at the end of the first stage are evaluated using Eqs. (3) and (22) respectively and salient results presented.

The increase in mass flow rate of deaerator water increases heat transfer coefficients in the first and second stages of the deaerator as shown in Fig. 2. The rate of increase is more in the first stage than in the second due to higher temperature potential between steam and water. A comparison of the heat transferred from both stages estimated from energy balance Eq. (8) with values estimated with Eq. (18) is in good agreement as can be seen from Fig. 3. This validates the heat transfer coefficients obtained with Eqs. (7) and (9) for the first and second stages respectively.

Experiments on direct contact condensation have been conducted with steamwater by Genic [10] and presented a regression equation between liquid phase transfer units and kinetic energy parameter given by

[NTU.sub.L.sup.C] = 0.185 [F.sup.-1.48.sub.LV] [28]

A comparison of the present data estimated for these parameters is in good agreement shown as Fig.4 valid for the deaerator first stage. The overall condensation heat transfer coefficient of two stage jet cum tray type deaerator is found to vary between 400-1600 W/([m.sup.2]K). Values estimated from the regression equation shown in Fig. 5 are in good agreement with values evaluated from energy balance Eq. (8). The increasing trends of mass transfer coefficient shown in Fig.6 for both the first and second stages of deaerator with increase in mass flow rate of water are similar to increases in heat transfer coefficient as can be seen from a comparison with Fig.2. The increase in deaerator pressure from 0.12 MPa to 0.2 MPa enhances the mass transfer coefficient in the second stage. This may be attributed to squeezing of oxygen from the water droplets due to higher pressure. The variation of Reynolds number with [Sh.sub.J] x [Sc.sup.-0.5.sub.J] for different operating conditions in the first and second stage of jet cum tray type deaerator is shown in Fig.7. The experimental values are in good agreement with the values estimated with the equation of Nosoko et al. [3] and Bakopoulos [11] for the first and second stages respectively. A plot between Sherwood number estimated with the regression equation is in good agreement with the experimental values as shown in Fig. 8 bringing out the validity of Eq. (27) proposed.

A comparison of the experimental values of oxygen stripped with that estimated with Eq. (22) for jet flow and with Eq. (26) for droplets is shown in Fig.9. A comparison of the values from theory and experiments is in good agreement.

[FIGURE 2 OMITTED]

[FIGURE 3 OMITTED]

[FIGURE 4 OMITTED]

[FIGURE 5 OMITTED]

[FIGURE 6 OMITTED]

[FIGURE 7 OMITTED]

[FIGURE 8 OMITTED]

[FIGURE 9 OMITTED]

Conclusions

The following conclusions can be drawn from the analysis of jet cum tray type deaerator

(a) The increase in mass flow rate of water increases the heat transfer coefficients whereas increase in pressure has negligible effect on the heat transfer coefficient for both stages of the deaerator.

(b) The increase in length of the second stage has no significant effect on heat transfer coefficients.

(c) The equation of Hasson et al.[1] is found to be valid for jet flow in the first stage only. The jet diameter is also an important parameter in determining the applicability of a theoretical equation in the estimation of heat transfer coefficient.

(d) The increase in flow rate of deaerator water increases the mass transfer coefficients in the first and second stages.

(e) The increase in deaerator pressure from 0.12 MPa to 0.2 MPa enhances the mass transfer coefficients in the second stage by 14 percent.

(f) The empirical correlations of Mayinger [8] predict the Sherwood number satisfactorily for both the first and second stages.

(g) The increase in length of the second stage has no significant effect on mass transfer coefficients.

(h) Regression Eqs. (19) and (27) for the estimation of heat and mass transfer coefficients are presented for use in deaerator design.

Nomenclature A surface area for heat transfer, [m.sup.2] C parameter defined in Eq. (9), g [D.sub.2.sup.3]/8[v.sup.2.sub.2] [C.sub.A] Concentration of oxygen, (= [[rho].sub.L] X) kg/[m.sup.3] [C.sub.P] specific heat, J/kg K d droplet diameter D inside diameter of the pipe/rivet sleeve, m [D.sub.L] diffusion coefficient or diffusivity of oxygen in water, [m.sup.2]/s [f.sub.1][f.sub.2][f.sub.3] parameters defined in Eq.(9) [F.sub.LV] kinetic energy parameter, 1/s[square root of ([[rho].sub.V] [[rho].sub.L])] g local acceleration of gravity, m/[s.sup.2] [Gr.sub.d] Grashof number, g [d.sup.3] ([[rho].sub.L] - [[rho].sub.V])/ [[rho].sub.L] [([[rho].sub.L]/ [[mu].sub.L])]).sup.2] Gz Graetz number h heat transfer coefficient, W/[m.sup.2]K [H.sub.fg] latent heat of condensation, J/kg Ja Jacob number, [C.sub.PL]([T.sub.s] - [T.sub.J])/[H.sub.fg] k thermal conductivity, W/m K [k.sub.L] mass transfer coefficient, m/s L length, m [L.sub.hb] hydraulic break-up length of the jet, m m mass flow rate of water, kg/s N number of droplets [NTU.sub.L] number of liquid phase transfer units, [T.sub.s] - [T.sub.i]/ [T.sub.S] - [T.sub.J] Nu Nusselt number [P.sub.D] deaerator pressure, N/[m.sup.2] ppb parts per billion Pr Prandtl number of liquid Jet in stage 2, [[upsilon].sub.2][[alpha].sub.2] Q heat transfer, W r radius of the liquid jet, m [Re.sub.d] droplet Reynolds number for mass transfer, d [V.sub.2]/[[upsilon].sub.V] [Re.sub.J] Jet Reynolds number, [GAMMA]/ [[upsilon].sub.L] [Re.sub.H] Reynolds number of liquid Jet in second stage, [V.sub.2][D.sub.2]/ 2[[upsilon].sub.2] s mass flow rate ratio of stripping steam to the inlet water Sc Schmidt number, [[upsilon].sub.L]/ [D.sub.L] [Sh.sub.0] initial Sherwood number [Sh.sub.d] droplet Sherwood number, [k.sub.L,d]d/[D.sub.L] [Sh.sub.J] Sherwood number of the liquid jet, [k.sub.L,J][delta]/[D.sub.L] St Stanton number t residence time, s T temperature, K [T.sub.AV] fluid properties evaluated at the average temperature, {([T.sub.o] + [T.sub.i]}/2}, K U overall heat transfer coefficient, W/[m.sup.2]K v volume of liquid, [m.sup.3] [v.sub.A] molecular volume of oxygen, [m.sup.3]/kg-mol V Velocity of liquid, m/s We Weber number of liquid Jet in second stage, {[[rho].sub.2][V.sup.2.sub.2] [D.sub.2]/2[sigma]} X mass fraction of oxygen in water, ppb z coordinate along the jet Greek symbols [alpha] thermal diffusivity, [m.sup.2]/s [GAMMA] volumetric flow rate per jet perimeter, [m.sup.2]/s [delta] Nusselt film thickness, [(3 [[upsilon].sub.L][GAMMA]/g).sup.1/3], m [theta] non-dimensional temperature at the end of the first stage [upsilon] kinematic viscosity, [m.sup.2]/s [mu] viscosity, kg/m s [rho] density, kg/[m.sup.3] [sigma] surface tension, {60.3 - 0.166 ([T.sub.J] - 273.15)[10.sup.-3]} N/m Subscripts AV average E energy balance EXP experiment d droplet H heat transfer i inlet J jet L liquid o outlet Reg regression equation S saturation t tray T theoretical V water vapor 1 first stage 2 second stage

Reference

[1] Hasson D., Luss D. and Peck R., 1964, "Theoretical analyses of vapor condensation on laminar liquid jets", Int. J. Heat and Mass transfer, vol.7, pp.969-981.

[2] Hasson D, Luss D, Navon U., 1964, "An experimental study of steam condensation on a laminar water sheet", Int. J. Heat and Mass transfer, vol.7, pp.983-1001.

[3] Nosoko T., Miyara A., Nagata T., 2002, " Characteristics of falling film flow on completely wetted horizontal tubes and the associated gas absorption", Int. Journal of Heat and Mass Transfer, Vol. 45, pp 2729-2738.

[4] Mochalova N.S., Kholpanov L.P., Malyusov V.A., 1988, "Heat transfer in vapor condensation on laminar and turbulent liquid jets, taking account of the inlet section and variability of the flow rate over the jet cross section", Institute of New Chemical problems, Academy of Sciences of the USSR, Moscow. Vol. 54, No. 5, pp. 732-735.

[5] Mills A.F., Kim S., Leininger T., Ofer S. and Pessran A., 1982, "Heat and Mass Transport in Turbulent Liquid Jets", International Journal of Heat and Mass Transfer, Vol.25,No.6, pp.889-897.

[6] Celata G.P, Cumo M, Farello G.E, and Focardi G., 1989, "A comprehensive analysis of direct contact condensation of saturated steam on subcooled liquid jets", Int. J. Heat Mass Transfer, vol.32, No.4, pp.639-654

[7] Hinze J.O., 1955, "Fundamentals of the hydrodynamic mechanism of splitting in dispersion processes" AIChEJ, pp 289-295.

[8] Mayinger, F., 1982, "Stromung und Warmeubergang in Flussigkeitsgemischen", Springer Verlag, Berlin

[9] Steinberger R.L., and Treybal R.E., 1960, "Mass transfer from a solid soluble sphere to a flowing liquid stream", A.I.Ch.E. J., vol.6, pp.227-232.

[10] Genic' S.B., 2006, "Direct-contact condensation heat transfer on downcomerless trays for steam-water system", Int. J. Heat Mass Transfer, vol.49, pp.1225-1230

[11] Bakopoulos A., 1980, "Liquid side controlled mass transfer in wetted-wall tube" German Chemical Engineering, Vol.3, pp. 241-252

[12] Emmert R.E. and Pigford R.L. 1954, "Interfacial Resistance--A case study of gas absorption in falling liquid films" , Chemical Engineering Progress, vol.50, No.2, pp.87-93

(1) K.V. Suryanarayana *, (2) K. V. Sharma, (3) P.K. Sarma, (4) V. Dharma Rao and (5) D.M. Reddy Prasad

(1) Department of Chemical Engineering, Sri Venkateswara Engineering College, Suryapet-508 213

(2) Faculty of Mechanical Engineering, Universiti Malaysia Pahang, 26300 Gambang, Kuantan, Pahang DM, E-mail: kvsharmajntu@yahoo.com

(3) International Director, GITAM University, Visakhapatnam-530 045 E-mail: sarmapk@yahoo.com

(4) Department of Chemical Engineering, A.U College of Engineering, Visakhapatnam-530 002 E-mail: v.dharmarao@yahoo.com

(5) Faculty of Chemical Engineering, Universiti Malaysia Pahang, 26300 Gambang, Kuantan, Pahang DM E-mail: dmrprasad@gmail.com

* Corresponding author Email: kagita_surya@yahoo.com

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Author: | Suryanarayana, K.V.; Sharma, K.V.; Sarma, P.K.; Rao, V. Dharma; Prasad, D.M. Reddy |
---|---|

Publication: | International Journal of Applied Engineering Research |

Article Type: | Report |

Geographic Code: | 1USA |

Date: | Nov 1, 2009 |

Words: | 4286 |

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