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Removal of dissolved oxygen in ultrapure water production using a membrane reactor
Pergamon
Chemical En(4ineerin~tScience, Vol. 50, No. 22, pp. 3547 3556, 1995 Copyright (~) 1995 Elsevier Science Ltd Printed in Great Britain. All rights reserved 0009 2509/95 $9.50 + 0.00
0009-2509(95)00192-1
R E M O V A L O F D I S S O L V E D O X Y G E N IN U L T R A P U R E WATER PRODUCTION USING A MEMBRANE REACTOR K. LI *''~, IVY C H U A ~, W. J. NG ~ and W. K. TEO ++ Department of Chemical Engineering and ~Department of Civil Engineering, National University of Singapore, l0 Kent Ridge Crescent, Singapore 0511 (First received 6 December 1994; revised manuscript received 18 May 1995: accepted 25 May 1995)
Abstract- Dissolved oxygen in water at parts per million (ppm) concentration range was reduced to a level less than 1.5 parts per billion (ppb) using a novel membrane reactor. The membrane reactor used was a polypropylene microporous hollow fibre membrane module packed with a palladium catalyst in the void space of the shell side. The saturated oxygen in water flowing in the shell side of the reactor was removed by purified hydrogen flowing in the fibre lumen. The hydrogen gas acted as both a reducing agent and a purge gas. Experimental results reveal that the rate of dissolved oxygen removal for both physical stripping and chemical reaction was controlled by the liquid film adjacent to the hollow fibre membrane and catalyst particles in the shell side. The removal of the dissolved oxygen was achieved by both the physical stripping and the chemical reaction at low catalyst loadings, whereas the chemical reaction became the dominant step at high catalyst loadings. A mass transfer correlation developed in this paper may be used in conjunction with available correlations for the design of a membrane deoxygenation reactor in the production of ultrapure water.
INTRODUCTION The production of ultrapure water is one of the key support services for the semiconductor, pharmaceutical, biotechnology, power and specialized chemical industries. Removal of dissolved oxygen from water is an important step in this production process and can be achieved by either physical or chemical methods. Conventional physical methods such as thermal degassing, vacuum degassing or nitrogen bubble deaeration have inherent drawbacks in terms of both operating costs and bulky construction. Also, with these physical methods, it is difficult to reduce the dissolved oxygen concentration from parts per million (ppm) level down to a few parts per billion level (ppb) (Kasama et al., 1990; Imaoka et al., 1991; Sato et al., 1991) which is often required in the semiconductor industry for wafer cleaning. The conventional chemical methods such as addition of hydrazine or sodium sulphite, although they provide an alternative to the physical methods, are undesirable because of the toxicity of the hydrazine material and because the addition of the sodium sulphite will result in an increase of the solid content of the water. Reaction with a reducing agent such as hydrogen in the presence of a catalyst to form water is an attractive method as it produces no by-product to contaminate the water, Kasama et al. (1990) have conducted a comprehensive study on removal of the dissolved oxygen from water. Four different systems have been considered in their study, namely nitrogen gas bubbling deaeration,
+Corresponding author,
catalytic reaction and hybrid systems of both these methods where a hollow fibre membrane module with oxygen selective membranes was utilized as degassing pretreatment prior to the nitrogen gas bubbling deaeration system and the catalytic reduction reaction, respectively. The experimental results obtained by them indicate that their second hybrid system gave the best deoxygenation results" and the DO levels between 3 and 5 ppb have been achieved. Tai et al. (1994) also studied the removal of the dissolved oxygen from water using membrane modules. Unlike the modules used by Kasama et al. (1990) where the membrane is non-porous oxygen selective type, the modules used by Tai et al. contain a hydrophobic microporous membrane. Water containing saturated oxygen was fed into the fibre lumen, while the purified nitrogen acting as purge gas was introduced into the shell side. Therefore the removal of dissolved oxygen from water is achieved by fast mass transfer rather than selective permeation. The experimental results obtained by them indicate that the hydrophobic membrane modules were capable of reducing the dissolved oxygen content in water to a level of around 8 ppb. The objective of this study was to investigate the removal of the dissolved oxygen from water using a membrane reactor where both physical and chemical processes take place simultaneously. The membrane reactor was made from a microporous hollow fibre module packed with a catalyst in the void space. Experiments have been carried out in this reactor with hydrogen gas used as both a purging gas as well as a reducing agent for the reduction of dissolved oxygen. This paper focuses on the effect of operating
3547
3548
K. LI et al.
conditions on the extent of dissolved oxygen removal and provides an experimental correlation which can be used for design of the reactor proposed,
the catalyst packed around the hollow fibres. Because the hollow fibre membrane used is highly hydrophobic, it has high resistance to water, but offers little resistance to gases. Therefore, the dissolved oxygen in the water diffuses through the hollow fibre membrane tO the gas stream and is purged away by the purified hydrogen flowing in the fibre lumen. Simultaneously, the hydrogen gas also diffuses through the hollow fibre membrane and dissolves into the water. Because of the presence of the catalyst, the dissolved hydrogen reacts with the remaining dissolved oxygen to form water which is not a subsequent contaminant. The purified hydrogen employed here not only acts as a purging gas (physical stripping), but also serves as a reducing agent (chemical reaction). Due to the simultaneous functions of both the physical and chemical methods in this membrane reactor, the dissolved oxygen level in the water at the outlet stream of the reactor may be much reduced. Oxygen concentration in the gas phase was measured by a Perkin Elmer (Autosystem) gas chromatograph using a thermal conductivity detector (TCD). The detector response was recorded and analysed by a personal computer with an integration package of "Omega" developed by Perkin Elmer. The results obtained together with the DO measurements from the water stream were used for the determination of oxygen balances and the DO removal by means of the physical stripping. An Orbisphere Model 3600 dissolved oxygen analyzer, from Orbisphere Laboratories, Switzerland, was used to measure the dissolved oxygen concentration in the feed and outlet streams. A sample flow of over 1 cm3/s through the analyser was found to be sufficient to overcome mixing problems in a stainless steel sample chamber. The accuracy of the analyser in
EXPERIMENTAL Hollow fibre membranes used in the present study were microporous polypropylene, of about 0.04 cm O D and 0.003 cm wall thickness, and were manufactured by Hoechst Celanese, U.S.A. The pores, about 30 nm in diameter, cover about 33% of the hollow fibre surface. The catalysts, obtained from Bayer in Germany, are palladium-doped anion exchange resins based on styrene and divinyl benzene. They are transparent-gelular and opaque-microporous spherical beads. The doping process results in the palladium being dispersed in the outer periphery of the beads as extremely fine particles in metallic form. Therefore, the dissolved hydrogen can rapidly access the active sites and the catalyst can attain a very high degree of efficiency, A membrane reactor fabricated using the abovementioned hollow fibre membrane and catalyst is shown in Fig. I(A). The hollow fibres were arranged either uniformly or randomly in the shell as shown in the transversal cross-sectional view of Fig. I(B) and (C), respectively. The catalyst particles were packed in the void space and were in close contact with the hollow fibres in the shell side. Four membrane reactors were used in this study. The characteristics of these reactors are given in Table 1. Figure 2 shows the experimental set-up used for the removal of the dissolved oxygen from water. Purified hydrogen is introduced into the fibre lumens, while the deionized water saturated with air (DO level of 8.3 ppm) is fed into the shell side and flows through
Waterout]-I
/
Catalyst
A
/hell
Hydrogenin~
~ ~Hollowfibres /
B
Waterin~TL
C Fig. 1. Schematic diagram of a membrane reactor.
Hydrogen out
Removal of dissolved oxygen in ultrapure water production
3549
Table 1. Characteristics of hollow fibre reactors Membrane reactor number
Number of fibres Length (m) Effective length (m) Potted length (m) Surface area of fibres (m 2) Packing fraction of fibres
l (uniform)
2 (random)
100 0.50 0.46 0.04 0.0577 0.04
500 050 0.46 0.04 0.2889 0.20
3 and 4 (random) 100 0.75 0.71 0.04 0.0891 0.04
Porosity of catalyst bed equals 0.63 (Chua, 1994) and has been used for the calculation of the surface area of the catalyst bead, A s . Packing fraction ~ = nd2o/d2, where d, is inside diameter of shell. Surface area of fibres, A,. = n x do × Zn.
Pressuregauge
?
Catalyst ~ =
(~
R°tametIer
~
Watereservi r or
confirmed with water completely deoxygenated by the addition of excess sodium sulphite with a catalyst of copper sulphate. The analyser requires only one point calibration, The measured dissolved oxygen data in both the inlet and outlet streams of the membrane reactors using the above-mentioned analyzer were used for the calculation of the overall mass transfer coefficient, KB:
Qdo~n( n
1 ) ] - QL/QgHn
.,,
Todrain
Fig2.. Experimentapparat al us.
the range of 0.1-8 ppm was l % or + 0 . 1 p p b ( w h i c h ever is greater). The response to changes in dissolved oxygen concentration took less than 10 min. Calibration of the oxygen analyser was achieved with water saturated with air at ambient conditions and
Kn
Vent
Rotameter
==
1-
t -~
[
C~Ut
1
× In (1 - QL/Q~HB) ci~" + (QL/QgHs)C~ ut) (1) Equation (1) was derived based on the countercurrent flow pattern shown in Fig. 3. A similar derivation of eq. (1) has been given by Sirkar (1992). It should be noted that in eq. (1), the overall mass transfer coefficient, Kn, is an overall value and is related to the coefficient of the overall mass transfer due to the physical stripping, KBL, and the coefficient of the overall mass transfer due to the chemical reaction, Ks.
Therelationofthesecoefficients,
Kn, K s L a n d K s , is
given in the next section of this paper.
3550
K. LI et aL
l Water out
QL C~a '
Water in
Liquidphase QL' ~
Gas phase boundary layer ~undary layer Membrane/ Liquid-solid interface
/
Q; p; H
c~ i| ~ ~Q"
Hydrogen o~ Hydrogen
.
¢~ ~
Catalyst
Pa Fig. 3. Countercurrent flow for the membrane reactor. Fig. 4. Concentration profile of dissolved oxygen in the membrane reactor.
THEORY Concentration profiles of the dissolved oxygen in a membrane reactor are shown in Fig. 4. The rate of the oxygen removal due to the physical stripping, NB1, can be expressed as Ns~ = kaL (CB -- CBi)A~ = kBm(PBi- pB~)A~ :
kBg (PBm - -
excess of the stoichiometric requirement, the order of the reaction is approximately 1.0 with respect to the oxygen concentration; therefore, at steady state, the reaction rate can be written as
PB) Am
(7)
NB2 = ( -- rs) = kBR Css As ~l
(2)
= KBL(Ca -- C*)Am
where kaL, kBm and ksg are the individual mass transfer coefficients in the liquid film, membrane and gas film, respectively. Am is the surface area of membrane. The dissolved oxygen concentration in equilibrium with partial pressure of oxygen in the gas phase is given in terms of Henry's law:
where kaR is the reaction rate constant. Rearranging eqs (5) and (7) yields Na2 = KsCsAs~l
(8)
l 1 1 K-s = ks~ + k-~R"
(9)
where
(3)
The rates of oxygen transfer due to the physical
Based on the additive rule of mass transfer resistances, the overall mass transfer coefficient, KBL, in eq. (2) can
stripping and the chemical reaction are in parallel; therefore, the total rate of oxygen removal is
be expressed as
Ns = N s l + Ns2 = KBL(CB -- C~)Am "k KsCsAsrl
PB = HsC*.
1
1
1
KsL = ks---L+ ~
1 + Hsks-'----gg"
= K s ( C s - C~)A,~
(4)
The rate equations for the oxygen removal due to the catalytic reaction can also be derived based on Fig. 4. The oxygen transfer rate from the bulk liquid to the surface of the catalyst is given by Na2 =
kBsAs~l(CB - Css)
(5)
where kBs is the mass transfer coefficient from the liquid to the catalyst interface, As is the total surface area of the catalyst, and CBs is the oxygen concentration at the surface of the catalyst. ~/is the area factor defined as Ae r / = A--~ (6) where Ae is the effective catalyst surface area. The reaction between oxygen and hydrogen takes place at the surface of the catalyst. It has been shown by Suppiah et al. (1988) that when hydrogen is present in
(10)
where KsCsAs~l
Ks = KsL + ( C a - C~)A,,"
(11)
When purge gas, i.e. hydrogen gas, flow rate is very high, i.e. C* --* 0, eq. (11) reduces to (12) A,, where KB is the overall oxygen transfer coefficient which consists of both resistances of physical stripping and catalytic chemical reaction. As can be seen from eq. (12), under the same operating conditions, such as the gas and liquid flow rates, temperatures and pressures, the overall mass transfer coefficient due to the physical stripping, KsL, and the overall mass transfer coefficient due to the chemical reaction, Ks, can be found by plotting Ka vs the surface area ratio, KB = KsL + Ks~l A s
As~Am.
Removal of dissolved oxygen in ultrapure water production The above derivation is based on the assumption that the dissolved hydrogen in water is always in excess compared with that of oxygen. Because of the presence of catalyst in the water, which has a strong affinity for the dissolved hydrogen (Bayer AG, 1990), the transfer of hydrogen from gas phase to the active sites of the catalyst is extremely fast and the mass transfer resistance presented by the hydrogen transfer is negligible compared to that of the oxygen transfer, This is true as it has been demonstrated by Li et al. (1994) that the large variation of hydrogen concentration (from 100 to 30%) has resulted in no effect on the dissolved oxygen removal• RESULTS AND DISCUSSION Experiments were conducted using both wellpacked and random-packed membrane reactors to evaluate the influence of operating conditions on the performance of the membrane reactors. The flow rates of water containing saturated oxygen were varied to establish the effect on the dissolved oxygen removal• Water pressures were always operated slightly higher than that of the gas-phase side of the membrane reactor to maintain the bubbleless operation required for the optimal function of the catalyst as small bubbles in water will reduce the catalyst activity dramatically (Bayer AG, 1990)• The overall mass transfer coefficients, Ke, calculated from eq. (1) gave a linear relationship when plotted against surface area ratio, As/A,, [see eq. (12)]. Typical experimental results are shown in Fig. 5. The mass transfer coefficients due to the physical stripping, KBL, and due to the catalytic reaction, Ks, were extracted from the intercept and slope of the plot, respectively. The regression results of Fig. 5 are given
1.oE-, S}
"
f
,~
'
o
5.0E-5 o..o.
E o
.
.
.
.
'
.
.
'
.
.
.
age as the gas and liquid flows can be operated independently and operating constraints such as flooding and loading, usually found in a packed column, do not exist in the membrane reactor proposed. Table 2 also shows that the mass transfer coefficient for physical stripping, KBL, increases with increase of water flow rate• This implies that the liquid film resistance adjacent to the membrane phase controls the dissolved oxygen stripping• The correlations for predieting the shell-side mass transfer of a membrane module alone have been developed by several investigators (Yang and Cussler, 1986; Prasad and Sirkar, 1988; Ahmed and Semmens, 1992; Costello et al., 1993; Li et al., 1994) shown in Table 3. Each of these
.
QL-4.17x10-6 m3/s .
.
.
-
,
"
0L=~.O0~10-6 ,,,3/, ,
o.o
'
OL,.a.33xlO_6m3/.
.
o.0
'
~
. 1"~°-'° s.oE-~
'-
.
l'°E-~"I
~t~ 5.0E-5
'
OL.2.50x10-6rn3/a
i
o.Q
' m3/,
in Table 2. As can be seen from the table, the values of intercepts are less than those of the slopes• The average contribution calculated based on these values of the mass transfer coefficients indicates that more than 90% of oxygen removal is achieved by catalytic reaction, while only about 10%o is achieved by the physical stripping process if the area ratio, As/A,,, is around 2. When the area ratio, As~Am, is considerably higher, the removal of the dissolved oxygen by physical stripping becomes small and negligible and the hollow fibre membrane used in this reactor becomes, somewhat, a gas distributor to ensure the continuous supply of the reducing agent, i.e. hydrogen gas. It has been shown by Imaoka et al. (1991) and Sato et al. (1991) that the chemical reduction of the dissolved oxygen is usually more effective than that of the physical method; therefore, a high area ratio, •4s/A,,. • may be selected for design purposes. However, the membrane area may also be an important factor as it directly relates to the hydrogen supply and its distribution within the reactor• Furthermore, use of membranes in this reactor provides an additional advant-
proposed correlations can be selectively employed for calculation of the dissolved oxygen removal by physical stripping depending on the packing densities of the hollow fibre membrane, water flow rate and the packing density of the catalyst in the shell side. For a low catalyst packing density or a partially expanded catalyst bed, the correlations mentioned above give an adequate estimation of shell-side mass transfer coefficient, Kec. When the packing density is high or for a fixed catalyst bed, the dissolved oxygen removed by physical stripping is much limited. A simple experiment using nitrogenwithinstead of hydrogen for theofmem-the brane reactor a high packing density
~
•0_ ~.0t-~| a.0~-5 ~
'
QL=l.67xI0-6
3551
.
i 0.3
.
,
.
.
.
o.e
Surface area r a t i o ,
,
.
i
o.g
,
la
As/Am
Fig, 5. Plot for determining mass transfer coefficients, K B t " and K s.
Table 2. Regression results of Fig. 5 Water flow QL (cmS/min)
Ke1~× 106 (m/s)
*Ks × 10~ (m/st
R2 (%)
100 150 200 250 300
2.59 3,38 4.37 5.30 6.25
3.41 4.07 4.27 4,5 l 4.73
98.3 97.3 97.2 96.6 96.1
*Ks values are calculated at q = 0.7.
K. LI et al.
3552
Table 3. Shell-side mass transfer correlations Reference Yang and Cussler (1986) Prasad and Sirkar (1988, Ahmed and Semmens (1992) Costello et al. (1993)
Lietal.(1994)
Range of Reynolds number
Correlations
(
Sh = 1.25 Re
Sh = 5.8(1- ~p)(~)Re°6Sc
Sh = 0.0104 Re°'s°6 Sh
0.1 < Re
SC TM
< 1000
0 < Re < 500
TM
500 < Re < 15,000
S C 0'33
= (0.53~0.581~)Re°'53Sc T M
-
/ d ",~o.967 Sh= 1.164~Rez) Sc °'33
-
10 < Re < I000
de = (4 × flow area)/(total fibre circumference).
catalyst indicates that the dissolved oxygen removal is much reduced. This is understandable as the high packing density of a catalyst tends to reduce the effective area of membrane exposed to water, and hence reduces the dissolved oxygen removed by means of physical stripping. As mentioned in the preceding paragraph, when the area ratio, As~Am, is high, contribution of dissolved oxygen removal by physical stripping becomes unimportant, and the dissolved oxygen removal is mainly achieved by the chemical reaction. Figure 6 shows the effect of water flow rate on the mass transfer coefficient for catalytic reaction using the area factor, ~/, as a parameter. The values of Ks are plotted vs water flow rate using log-log co-ordinates. The hydrogen flow rate in the fibre lumen was maintained at a constant rate of 401/h. As would be expected, for a given water flow rate, the mass transfer coefficient, Ks, is decreased with an increase of area factor, ~/, The regression analysis shows that the best fitted lines drawn through the data points correspond to the variation of Ks as 0.28 to 0.31 power of water flow rate. The increased mass transfer coefficient with water flow rate suggested that the removal of the dissolved oxygen due to the catalytic reaction in the membrane reactor is primarily controlled by the resistance at the liquid-solid boundary layer, i.e. I/ks. This conclusion drawn from the above experimental results is consistent with that of earlier observations (Tay, 1994) for the catalyst bed (randomly packed) reactor used for removal of the dissolved oxygen, The membrane reactors employed in the present study may be viewed as an expanded bed reactor where two different motion behaviours of the catalyst particles were observed in the reactor depending on the a m o u n t of the catalyst used and the water flow rate. These two behaviours are illustrated in Fig. 7. As can be seen from the figure, at a low volume of packed catalyst and a high water flow rate, the catalyst particles are disturbed so that expanded or jiggling bed behaviour is observed [Fig. 7(A)]. When the volume of the packed catalyst particles is high and water flow rate is low, the particles are not disturbed; in this case,
sOoox~o - s Module 1
10o0 = "~ b~ ~
_:
~. ' ° ' ~ ----L--~ ---
100
'r/-0.4
-
~
~
~-1.o
10
u u
~ u,du~,z ~ 10oo ~ .-.------.---loo ~
lo
,~-o.1
-
_
2
3
_
~-0.4 ~-'~i~
5
7.
10x10
-8
Woter flow rote, m 3 / s Fig. 6. Effect ofwater flow rate on the observed mass transfer coefficient, K s.
the membrane reactor acts like a fixed bed [Fig. 7(B)]. The value of the mass transfer coefficients in fluidized bed or fixed bed reactor is often expressed as dimensionless correlations. The appropriate correlations give the Sherwood number as a function of the Reynolds n u m b e r and Schmidt n u m b e r as shown:
Sh = A Re=Sc ~
(13)
or (~-~Z)
(vLdpy(v'~ = A \ - ~ - - / \~-~/
(14)
where A, ¢ and fl are adjustable parameters. Since, in this study, the Schmidt number, Sc, was not varied, a 0.33 power dependence, which is commonly accepted in the literature, was assumed. The values of A and ¢ were obtained by least-square analysis. The
Removal of dissolved oxygen in ultrapure water production
high fluid
~
3553
.,~ep' low fluid
expanded bed behaviour
A
13 Fig. 7. Catalyst particle motion in the membrane reactor.
normalized experimental data have been presented in different ways shown in Figs 8,9 and 10, and is discussed below. Figure 8 plots the mass transfer coefficient, Ks, as Sherwood number, Sh, vs Reynolds number, Re, in log-log form, while Fig. 9 shows a plot which takes account of the area factor, q, and membrane packing fraction, ~b, in the reactor. It can be seen from Fig. 8 that the reactors with either uniformly packed or randomly packed hollow fibres have a similar performance indicating that the way in which the hollow fibres are packed in the reactor is not important as long as the catalyst packed on the shell side surrounds the hollow fibres in the reactor. The general trends of the lines in Fig. 8 are similar to those correlations available in the literature for both fixed and fluidized bed reactors (Chu et al., 1953; Frantz, 1962; Upadhyay and Tripathi, 1975; Dwivedi and Upadhyay, 1977}. If the modified Sherwood, Sh' = {ksdp/DB)fl, and Reynolds, Re' = (vLdp/v) (1 4'), numbers are used, a more general dimensionless correlation can be obtained, and is shown in Fig. 9 where the regression results have been shown to be best described by -
-
Sh =
5.05
[Re(l -
~)]0.307
(15)
50o m ~
• Modore~.~=0.~ • Mod~2, ~=o.~
%~ o o ~ ~.
c II
~ ~g o ~
0
~
,
• Module2, ~=~.o
~
~o
20
~o
Reynolds number, (VLdp/V)
Fig. 8. The observed variation in Sherwood number, Sh, with Reynolds number, Re, for dissolved oxygen removal in the membrane reactors. -% .~. ~
/
,'F 1o i
~
Mod~,e~
]
Module2
" ~'-"-f-
-g. ~= ~ ~
t [
sh'~=5"os(Re) °'3°7
to~
/
R2=97'8~
~g
or
Modified Reynolds number, Re=(VLdp/V)(1-*)
Sh = 0.40 [Re(l -- (~1)]0"307 Sc O'33
(16) r/~ Equation (16) treats the void fraction as a constant value (under fixed bed condition), which may not be the case if the reactor with low catalyst packing densities is operated with a high water flow rate. As mentioned earlier, under certain operating conditions, the catalyst bed is expanded and this will inevitably increase the void fraction of the catalyst bed. The following linear relation may be used to correlate the catalyst bed expansion: Z 1 ~:i = el~-/,
~:i < 1
t17)
Fig. 9. The observed variation in modified Sherwood number, Sh', with modified Reynolds number, Re', for dissolved oxygen removal in the membrane reactors.
where e~ and Z~ are void fraction and length of the catalyst bed under fixed bed conditions, respectively, while e~ and Z~ are void fraction and length of the catalyst bed under expanded bed conditions, respectively. By incorporation ofe~ into the Reynolds humber, the modified Reynolds number, Re" is defined as Re" =
l!Ldp(1 -- q)) v(l - - el)
(18)
K. Ll et al.
3554
Plotting the modified Sherwood number, (ksdp/OB)q, vs the modified Reynolds number, Re", shown in Fig. 10, the regression results give 0.55 ~ R e ( 1 - q~)]°22 Sh = SC O'33 (19) L 1--- ~i _l '
Figure 11 shows the effect of water flow rate on the dissolved oxygen concentration in the product stream. The experimental data were obtained using reactors 3 and 4 connected in parallel. The total volumes of the catalyst packed in these two reactors are 400 cm3; thus, the area ratio, As/A,, is very high
The correlations in eqs(16) and (19) have similar trends to many other available correlations suggested earlier as shown in Table 4. The mass transfer coefficient Ks varies as a 0.22 power with Re", which is consistent with the 0.22 power (Chu et al,, 1953) and the 0.21 power (Upadhyay and Tripathi, 1975) found by previous workers. In comparison with Sh vs Re', the mass transfer coefficient, Ks, is found to vary with the 0.307 power with Re', which is slightly different from that of 0.28 given by Dwivedi and Upadhyay (1977). The numerical constants found here cannot be directly compared with those given in Table 4 as, in this study, the area factor, q, is an unknown value which dictates the final numerical constants. If the constants listed in Table 4 are assumed to be true, the area factor of the catalyst used in this study is around 10-40%.
and is equal to 10. Under such a condition, K B ~ K s , and eqs (1) and (15) can therefore be used for prediction purposes. The calculated data are shown by the solid curve. As can be seen from Fig. 11, the observed experimental data and the calculated results are in good agreement. The dissolved oxygen level in the outlet stream can be achieved as low as 1.2 ppb, which is basically impossible to achieve using conventional physical methods. CONCLUSIONS A novel membrane reactor has been developed for removal of the dissolved oxygen in water from parts per million (ppm) to a level less than 1.5 ppb. The membrane reactor comprises a microporous hollow fibre module packed with a palladium catalyst in the void space. Mass transfer studies show that the over-
~ o_
:~u~ ~* -~ "
• Module I • Module 2
lo
• Exp. data -Calculated values
•
~
R2=96.1% 1
~0
~ ~~ ~_ 50 ~ E 40
/
~ .E
10
~'~
0
4
Modified Reynolds number, R'e=(VLdp/#)(1-$)/(1-~i)
8 Water
Fig. 10. The observed variation in modified Sherwood num-
' 16
12
20x10
-6
mS/s
flow fete,
Fig. 11. Comparison between experimental results obtained
bar, Sh', with modified Reynolds number, Re", for dissolved fromreactors 3 and 4 and calculated results from eqs (1) and oxygen removal in the membrane reactors.
(15).
Table 4. Comparison of correlations Reference Chu et al. (1953) Frantz (1962)
Correlations
/ Re ~T M Sh = 5. ~l_e.i -Sc T M [ Re \ °'3°
Sh = 5 \11--1ei.}
Sc°33
Range of Reynolds number
1
< ( Re ~ \I - eiJ < 30
1
< ( Re ~ \1 - eJ < 80
(Re~
(Re) ~
Upadhyay and Tripathi (1975)
Sh = 3.8155k1-~/) T M Sc°'Sa
Dwivedi and Upadhyay (1977)
Sh = 1.1068 ReO.28SCO.33
Re < 10
This study
Sh = 0.40 ERe(1 -- q~)]o.3o7Sco.aa
Re(1 - ~b)< 10
= n k ~ J
\
l-e,
< 20
)
50
Removal of dissolved oxygen in ultrapure water production all mass transfer coefficient for physical stripping, KsL, is generally smaller than that for chemical reaction, Ks. Both KBL and Ks are dependent on the water flow rate indicating that the liquid films adjacent to the hollow fibre membrane and the solid catalyst surface control the removal of the dissolved oxygen. The mass transfer study further reveals that at high catalyst loadings, the total oxygen removal is mainly achieved by the chemical reaction alone and the removal of the dissolved oxygen by physical stripping is negligible. Under such a condition, the hollow fibre membrane employed becomes akin to a gas distributor and ensures a constant supply of purified hydrogen for the catalytic reaction. Thus the membranes provide a stationary interface which enables both the gas and liquid flows to be operated independently and operation constraints such as flooding and loading points, usually found in packed columns, no longer exist. Based on the experimental results, a correlation has been developed, which can be used for the design of membrane reactors of the type described in this paper. Under high catalyst loadings, the correlation proposed can be used alone for design purposes, whilst at low catalyst loadings, i.e. both chemical reaction and physical stripping are significant for the dissolved oxygen removal, the correlations listed in Table 3 and the correlation developed must be employed together for the design of the membrane reactor.
Acknowledgements--The authors are grateful to the Singapore Science Council for the award of RDAS Grant (ST/90/03) to support the research and development work on ultrapure water production. NOTATION A At A,~ As C8
constant in eqs (13) and (14) effective catalyst surface area, m 2 total membrane area, m 2 total surface area of catalyst, m 2 dissolved oxygen concentration, kmol/m 3
do dp HB k
outside diameter of hollow fibres, m catalyst particle diameter, m Henry's law constant individual mass transfer coefficient
kns KB
reaction rate constant, m/s overall mass transfer coefficient of oxygen defined in eq. (12), m/s overall mass transfer coefficient of oxygen for physical stripping, m/s
KnL Ks n NB
overall mass transfer coefficient of oxygen for catalytic reaction, m/s number of fibres rate of the dissolved oxygen transfer, kmol/s
Q Re Re' Re"
fluid flow rate, m3/s Reynolds number ( = vcdp/v) modified Reynolds number(= vcd~/v)(1 - ~b) modified Reynolds number ( = vLdp/v) (1 -q~)/(1 -- el)
Sh Sh'
Sherwood number ( = Ksdp/DB) modified Sherwood number (= Ksdp/DB)q
CE$ 50-27-E
vL Z Z1 Zi
3555
velocity of water in membrane reactor without catalyst packings, m/s effective length of the fibres, m length of catalyst bed at fixed bed condition, m length of catalyst bed at fluidized bed condition, m
Greek letters ct fl ¢ el
parameter in eqs (13) and (14) parameter in eqs (13) and (14) porosity of catalyst bed void fraction of fixed catalyst bed
e.i q v q~
void fraction of fluidized catalyst bed area factor, defined in eq. (6) kinematic viscosity, m2/s packing fraction of hollow fibres
Subscripts 1 physical stripping 2 catalytic reaction g gas i interface L liquid m membrane S catalyst surface Superscripts in reactor inlet out reactor outlet . liquid phase in equilibrium with gas phase
REFERENCES Ahmed, T. and Semmens, M. J., 1992, Use of sealed end hollow fibres for bubbleless membrane aeration: experimental studies. J. Membrane Sci. 69, 1-10. Bayer AG, 1990, Catalytic removal of dissolved oxygen from water. Organic Chemical Division, Forschung/Anwendungstechnik Lewatit, Ionenaustauscher/Katalysatoren, Geb. B106, Leverkusen, pp. 1~. Chu, J. C., Kalil, J. and Wetteroth, W. A., 1953, Mass transfer in a fluidized bed. Chem. En#ng Prog. 49, 141-149. Chua, I., 1994, Elimination of dissolved oxygen in ultrapure water by chemical reduction and physical permeation processes. M.Eng thesis, Department of Chemical Engineering, National University of Singapore. Costello,M. J., Fane, A. G., Hogan, P. A. and Schofield, R.W., 1993, The effect of shell side hydrodynamics on the performanceof hollow fibre modules. J. Membrane Sci. 80, 1-11. Dwivedi, P. N. and Upadhyay, S. N., 1977, Particle-fluid mass transfer in fixed and fluidizedbeds. Ind. Engn# Chem. Des. Dev. 16, 157-165. Frantz, J., 1962, Design for fluidization. Chem. Engn9 69, 161-178. Imaoka, T., Yagi, Y., Kasama, Y., Sugiyama, 1., Isagawa, T. and Ohmi, T., 1991,Advanced ultrapure water systemsfor ULSI processing, in Proceedinos of the lOth Annual SemiconductorPure Water Conference, Santa Clara, CA, 26-28 February (Edited by M. K. Balazs), pp. 128-146. Kasama, Y., Yagi, Y., Imaoka, T., Kawakami, M. and Ohmi, T., 1990, Advanced D.I. water system with low dissolved oxygen for ULSI processing. Proceedinos Institute of EnvironmentalScience, pp. 344-349.
3556
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Li, K., Chua, 1., Ng, W. J. and Teo, W. K., 1994, Membrane reactor for the removal of dissolved oxygen from water. U.K. Patent, App. No. 9321867.5. Li, K., Tai, M. S. L. and Teo, W. K., 1994, Design of a CO 2 scrubber for self-contained breathing systems using a microporous membrane. J. Membrane Sci. 86, 119-125. Prasad, R. and Sirkar, K. K., 1988, Dispersion-free solvent extraction with microporous hollow fibre modules. A.I.Ch.E.J. 34, 177-188. Sato, H., Hashimoto, N., Shinoda, T. and Takino, K., 1991, Dissolved oxygen removal in ultrapure water for semiconductor processes, in Proceedings of the lOth Annual Semiconductor Pure Water Conference, Santa Clara, CA, 26--28 February (Edited by M. K. Balazs), pp. 147-164. Sirkar, K. K., 1992, Other new membrane processes, in Membrane Handbook (Edited by W. S. Winston Ho and K. Sirkar), pp. 891-992. Van Nostrand Reinhold, New York, NY.
Suppiah, S., Kutchcoskie, K. J., Balakrishnan, P. V. and Chuang, K. T., 1988, Dissolved oxygen removal by combination with hydrogen using wetproofed catalysts. Can. J. Chem. Engng 66, 849-857. Tai, M. S. L., Chua, I., Li, K., Ng, W. J. and Teo, W. K., 1994, Removal of dissolved oxygen in ultrapure water production using microporous membrane modules. J. Membrane Sci. 87,99-105. Tay, Y. T., 1994, Removal of dissolved oxygen using random packed catalyst bed. Final year, B.Engng Research Project, Department of Chemical Engineering, National University of Singapore. Upadhyay, S. N. and Tripathi, G., 1975, Liquid phase mass transfer in fixed and fluidized beds of large particles. J. Chem. Engno Data 20, 20-26. Yang, M. and Cussler, E. L., 1986, Designing hollow-fibre contactors. A.I.Ch.E.J. 32, 1910-1916.
Chemical En(4ineerin~tScience, Vol. 50, No. 22, pp. 3547 3556, 1995 Copyright (~) 1995 Elsevier Science Ltd Printed in Great Britain. All rights reserved 0009 2509/95 $9.50 + 0.00
0009-2509(95)00192-1
R E M O V A L O F D I S S O L V E D O X Y G E N IN U L T R A P U R E WATER PRODUCTION USING A MEMBRANE REACTOR K. LI *''~, IVY C H U A ~, W. J. NG ~ and W. K. TEO ++ Department of Chemical Engineering and ~Department of Civil Engineering, National University of Singapore, l0 Kent Ridge Crescent, Singapore 0511 (First received 6 December 1994; revised manuscript received 18 May 1995: accepted 25 May 1995)
Abstract- Dissolved oxygen in water at parts per million (ppm) concentration range was reduced to a level less than 1.5 parts per billion (ppb) using a novel membrane reactor. The membrane reactor used was a polypropylene microporous hollow fibre membrane module packed with a palladium catalyst in the void space of the shell side. The saturated oxygen in water flowing in the shell side of the reactor was removed by purified hydrogen flowing in the fibre lumen. The hydrogen gas acted as both a reducing agent and a purge gas. Experimental results reveal that the rate of dissolved oxygen removal for both physical stripping and chemical reaction was controlled by the liquid film adjacent to the hollow fibre membrane and catalyst particles in the shell side. The removal of the dissolved oxygen was achieved by both the physical stripping and the chemical reaction at low catalyst loadings, whereas the chemical reaction became the dominant step at high catalyst loadings. A mass transfer correlation developed in this paper may be used in conjunction with available correlations for the design of a membrane deoxygenation reactor in the production of ultrapure water.
INTRODUCTION The production of ultrapure water is one of the key support services for the semiconductor, pharmaceutical, biotechnology, power and specialized chemical industries. Removal of dissolved oxygen from water is an important step in this production process and can be achieved by either physical or chemical methods. Conventional physical methods such as thermal degassing, vacuum degassing or nitrogen bubble deaeration have inherent drawbacks in terms of both operating costs and bulky construction. Also, with these physical methods, it is difficult to reduce the dissolved oxygen concentration from parts per million (ppm) level down to a few parts per billion level (ppb) (Kasama et al., 1990; Imaoka et al., 1991; Sato et al., 1991) which is often required in the semiconductor industry for wafer cleaning. The conventional chemical methods such as addition of hydrazine or sodium sulphite, although they provide an alternative to the physical methods, are undesirable because of the toxicity of the hydrazine material and because the addition of the sodium sulphite will result in an increase of the solid content of the water. Reaction with a reducing agent such as hydrogen in the presence of a catalyst to form water is an attractive method as it produces no by-product to contaminate the water, Kasama et al. (1990) have conducted a comprehensive study on removal of the dissolved oxygen from water. Four different systems have been considered in their study, namely nitrogen gas bubbling deaeration,
+Corresponding author,
catalytic reaction and hybrid systems of both these methods where a hollow fibre membrane module with oxygen selective membranes was utilized as degassing pretreatment prior to the nitrogen gas bubbling deaeration system and the catalytic reduction reaction, respectively. The experimental results obtained by them indicate that their second hybrid system gave the best deoxygenation results" and the DO levels between 3 and 5 ppb have been achieved. Tai et al. (1994) also studied the removal of the dissolved oxygen from water using membrane modules. Unlike the modules used by Kasama et al. (1990) where the membrane is non-porous oxygen selective type, the modules used by Tai et al. contain a hydrophobic microporous membrane. Water containing saturated oxygen was fed into the fibre lumen, while the purified nitrogen acting as purge gas was introduced into the shell side. Therefore the removal of dissolved oxygen from water is achieved by fast mass transfer rather than selective permeation. The experimental results obtained by them indicate that the hydrophobic membrane modules were capable of reducing the dissolved oxygen content in water to a level of around 8 ppb. The objective of this study was to investigate the removal of the dissolved oxygen from water using a membrane reactor where both physical and chemical processes take place simultaneously. The membrane reactor was made from a microporous hollow fibre module packed with a catalyst in the void space. Experiments have been carried out in this reactor with hydrogen gas used as both a purging gas as well as a reducing agent for the reduction of dissolved oxygen. This paper focuses on the effect of operating
3547
3548
K. LI et al.
conditions on the extent of dissolved oxygen removal and provides an experimental correlation which can be used for design of the reactor proposed,
the catalyst packed around the hollow fibres. Because the hollow fibre membrane used is highly hydrophobic, it has high resistance to water, but offers little resistance to gases. Therefore, the dissolved oxygen in the water diffuses through the hollow fibre membrane tO the gas stream and is purged away by the purified hydrogen flowing in the fibre lumen. Simultaneously, the hydrogen gas also diffuses through the hollow fibre membrane and dissolves into the water. Because of the presence of the catalyst, the dissolved hydrogen reacts with the remaining dissolved oxygen to form water which is not a subsequent contaminant. The purified hydrogen employed here not only acts as a purging gas (physical stripping), but also serves as a reducing agent (chemical reaction). Due to the simultaneous functions of both the physical and chemical methods in this membrane reactor, the dissolved oxygen level in the water at the outlet stream of the reactor may be much reduced. Oxygen concentration in the gas phase was measured by a Perkin Elmer (Autosystem) gas chromatograph using a thermal conductivity detector (TCD). The detector response was recorded and analysed by a personal computer with an integration package of "Omega" developed by Perkin Elmer. The results obtained together with the DO measurements from the water stream were used for the determination of oxygen balances and the DO removal by means of the physical stripping. An Orbisphere Model 3600 dissolved oxygen analyzer, from Orbisphere Laboratories, Switzerland, was used to measure the dissolved oxygen concentration in the feed and outlet streams. A sample flow of over 1 cm3/s through the analyser was found to be sufficient to overcome mixing problems in a stainless steel sample chamber. The accuracy of the analyser in
EXPERIMENTAL Hollow fibre membranes used in the present study were microporous polypropylene, of about 0.04 cm O D and 0.003 cm wall thickness, and were manufactured by Hoechst Celanese, U.S.A. The pores, about 30 nm in diameter, cover about 33% of the hollow fibre surface. The catalysts, obtained from Bayer in Germany, are palladium-doped anion exchange resins based on styrene and divinyl benzene. They are transparent-gelular and opaque-microporous spherical beads. The doping process results in the palladium being dispersed in the outer periphery of the beads as extremely fine particles in metallic form. Therefore, the dissolved hydrogen can rapidly access the active sites and the catalyst can attain a very high degree of efficiency, A membrane reactor fabricated using the abovementioned hollow fibre membrane and catalyst is shown in Fig. I(A). The hollow fibres were arranged either uniformly or randomly in the shell as shown in the transversal cross-sectional view of Fig. I(B) and (C), respectively. The catalyst particles were packed in the void space and were in close contact with the hollow fibres in the shell side. Four membrane reactors were used in this study. The characteristics of these reactors are given in Table 1. Figure 2 shows the experimental set-up used for the removal of the dissolved oxygen from water. Purified hydrogen is introduced into the fibre lumens, while the deionized water saturated with air (DO level of 8.3 ppm) is fed into the shell side and flows through
Waterout]-I
/
Catalyst
A
/hell
Hydrogenin~
~ ~Hollowfibres /
B
Waterin~TL
C Fig. 1. Schematic diagram of a membrane reactor.
Hydrogen out
Removal of dissolved oxygen in ultrapure water production
3549
Table 1. Characteristics of hollow fibre reactors Membrane reactor number
Number of fibres Length (m) Effective length (m) Potted length (m) Surface area of fibres (m 2) Packing fraction of fibres
l (uniform)
2 (random)
100 0.50 0.46 0.04 0.0577 0.04
500 050 0.46 0.04 0.2889 0.20
3 and 4 (random) 100 0.75 0.71 0.04 0.0891 0.04
Porosity of catalyst bed equals 0.63 (Chua, 1994) and has been used for the calculation of the surface area of the catalyst bead, A s . Packing fraction ~ = nd2o/d2, where d, is inside diameter of shell. Surface area of fibres, A,. = n x do × Zn.
Pressuregauge
?
Catalyst ~ =
(~
R°tametIer
~
Watereservi r or
confirmed with water completely deoxygenated by the addition of excess sodium sulphite with a catalyst of copper sulphate. The analyser requires only one point calibration, The measured dissolved oxygen data in both the inlet and outlet streams of the membrane reactors using the above-mentioned analyzer were used for the calculation of the overall mass transfer coefficient, KB:
Qdo~n( n
1 ) ] - QL/QgHn
.,,
Todrain
Fig2.. Experimentapparat al us.
the range of 0.1-8 ppm was l % or + 0 . 1 p p b ( w h i c h ever is greater). The response to changes in dissolved oxygen concentration took less than 10 min. Calibration of the oxygen analyser was achieved with water saturated with air at ambient conditions and
Kn
Vent
Rotameter
==
1-
t -~
[
C~Ut
1
× In (1 - QL/Q~HB) ci~" + (QL/QgHs)C~ ut) (1) Equation (1) was derived based on the countercurrent flow pattern shown in Fig. 3. A similar derivation of eq. (1) has been given by Sirkar (1992). It should be noted that in eq. (1), the overall mass transfer coefficient, Kn, is an overall value and is related to the coefficient of the overall mass transfer due to the physical stripping, KBL, and the coefficient of the overall mass transfer due to the chemical reaction, Ks.
Therelationofthesecoefficients,
Kn, K s L a n d K s , is
given in the next section of this paper.
3550
K. LI et aL
l Water out
QL C~a '
Water in
Liquidphase QL' ~
Gas phase boundary layer ~undary layer Membrane/ Liquid-solid interface
/
Q; p; H
c~ i| ~ ~Q"
Hydrogen o~ Hydrogen
.
¢~ ~
Catalyst
Pa Fig. 3. Countercurrent flow for the membrane reactor. Fig. 4. Concentration profile of dissolved oxygen in the membrane reactor.
THEORY Concentration profiles of the dissolved oxygen in a membrane reactor are shown in Fig. 4. The rate of the oxygen removal due to the physical stripping, NB1, can be expressed as Ns~ = kaL (CB -- CBi)A~ = kBm(PBi- pB~)A~ :
kBg (PBm - -
excess of the stoichiometric requirement, the order of the reaction is approximately 1.0 with respect to the oxygen concentration; therefore, at steady state, the reaction rate can be written as
PB) Am
(7)
NB2 = ( -- rs) = kBR Css As ~l
(2)
= KBL(Ca -- C*)Am
where kaL, kBm and ksg are the individual mass transfer coefficients in the liquid film, membrane and gas film, respectively. Am is the surface area of membrane. The dissolved oxygen concentration in equilibrium with partial pressure of oxygen in the gas phase is given in terms of Henry's law:
where kaR is the reaction rate constant. Rearranging eqs (5) and (7) yields Na2 = KsCsAs~l
(8)
l 1 1 K-s = ks~ + k-~R"
(9)
where
(3)
The rates of oxygen transfer due to the physical
Based on the additive rule of mass transfer resistances, the overall mass transfer coefficient, KBL, in eq. (2) can
stripping and the chemical reaction are in parallel; therefore, the total rate of oxygen removal is
be expressed as
Ns = N s l + Ns2 = KBL(CB -- C~)Am "k KsCsAsrl
PB = HsC*.
1
1
1
KsL = ks---L+ ~
1 + Hsks-'----gg"
= K s ( C s - C~)A,~
(4)
The rate equations for the oxygen removal due to the catalytic reaction can also be derived based on Fig. 4. The oxygen transfer rate from the bulk liquid to the surface of the catalyst is given by Na2 =
kBsAs~l(CB - Css)
(5)
where kBs is the mass transfer coefficient from the liquid to the catalyst interface, As is the total surface area of the catalyst, and CBs is the oxygen concentration at the surface of the catalyst. ~/is the area factor defined as Ae r / = A--~ (6) where Ae is the effective catalyst surface area. The reaction between oxygen and hydrogen takes place at the surface of the catalyst. It has been shown by Suppiah et al. (1988) that when hydrogen is present in
(10)
where KsCsAs~l
Ks = KsL + ( C a - C~)A,,"
(11)
When purge gas, i.e. hydrogen gas, flow rate is very high, i.e. C* --* 0, eq. (11) reduces to (12) A,, where KB is the overall oxygen transfer coefficient which consists of both resistances of physical stripping and catalytic chemical reaction. As can be seen from eq. (12), under the same operating conditions, such as the gas and liquid flow rates, temperatures and pressures, the overall mass transfer coefficient due to the physical stripping, KsL, and the overall mass transfer coefficient due to the chemical reaction, Ks, can be found by plotting Ka vs the surface area ratio, KB = KsL + Ks~l A s
As~Am.
Removal of dissolved oxygen in ultrapure water production The above derivation is based on the assumption that the dissolved hydrogen in water is always in excess compared with that of oxygen. Because of the presence of catalyst in the water, which has a strong affinity for the dissolved hydrogen (Bayer AG, 1990), the transfer of hydrogen from gas phase to the active sites of the catalyst is extremely fast and the mass transfer resistance presented by the hydrogen transfer is negligible compared to that of the oxygen transfer, This is true as it has been demonstrated by Li et al. (1994) that the large variation of hydrogen concentration (from 100 to 30%) has resulted in no effect on the dissolved oxygen removal• RESULTS AND DISCUSSION Experiments were conducted using both wellpacked and random-packed membrane reactors to evaluate the influence of operating conditions on the performance of the membrane reactors. The flow rates of water containing saturated oxygen were varied to establish the effect on the dissolved oxygen removal• Water pressures were always operated slightly higher than that of the gas-phase side of the membrane reactor to maintain the bubbleless operation required for the optimal function of the catalyst as small bubbles in water will reduce the catalyst activity dramatically (Bayer AG, 1990)• The overall mass transfer coefficients, Ke, calculated from eq. (1) gave a linear relationship when plotted against surface area ratio, As/A,, [see eq. (12)]. Typical experimental results are shown in Fig. 5. The mass transfer coefficients due to the physical stripping, KBL, and due to the catalytic reaction, Ks, were extracted from the intercept and slope of the plot, respectively. The regression results of Fig. 5 are given
1.oE-, S}
"
f
,~
'
o
5.0E-5 o..o.
E o
.
.
.
.
'
.
.
'
.
.
.
age as the gas and liquid flows can be operated independently and operating constraints such as flooding and loading, usually found in a packed column, do not exist in the membrane reactor proposed. Table 2 also shows that the mass transfer coefficient for physical stripping, KBL, increases with increase of water flow rate• This implies that the liquid film resistance adjacent to the membrane phase controls the dissolved oxygen stripping• The correlations for predieting the shell-side mass transfer of a membrane module alone have been developed by several investigators (Yang and Cussler, 1986; Prasad and Sirkar, 1988; Ahmed and Semmens, 1992; Costello et al., 1993; Li et al., 1994) shown in Table 3. Each of these
.
QL-4.17x10-6 m3/s .
.
.
-
,
"
0L=~.O0~10-6 ,,,3/, ,
o.o
'
OL,.a.33xlO_6m3/.
.
o.0
'
~
. 1"~°-'° s.oE-~
'-
.
l'°E-~"I
~t~ 5.0E-5
'
OL.2.50x10-6rn3/a
i
o.Q
' m3/,
in Table 2. As can be seen from the table, the values of intercepts are less than those of the slopes• The average contribution calculated based on these values of the mass transfer coefficients indicates that more than 90% of oxygen removal is achieved by catalytic reaction, while only about 10%o is achieved by the physical stripping process if the area ratio, As/A,,, is around 2. When the area ratio, As~Am, is considerably higher, the removal of the dissolved oxygen by physical stripping becomes small and negligible and the hollow fibre membrane used in this reactor becomes, somewhat, a gas distributor to ensure the continuous supply of the reducing agent, i.e. hydrogen gas. It has been shown by Imaoka et al. (1991) and Sato et al. (1991) that the chemical reduction of the dissolved oxygen is usually more effective than that of the physical method; therefore, a high area ratio, •4s/A,,. • may be selected for design purposes. However, the membrane area may also be an important factor as it directly relates to the hydrogen supply and its distribution within the reactor• Furthermore, use of membranes in this reactor provides an additional advant-
proposed correlations can be selectively employed for calculation of the dissolved oxygen removal by physical stripping depending on the packing densities of the hollow fibre membrane, water flow rate and the packing density of the catalyst in the shell side. For a low catalyst packing density or a partially expanded catalyst bed, the correlations mentioned above give an adequate estimation of shell-side mass transfer coefficient, Kec. When the packing density is high or for a fixed catalyst bed, the dissolved oxygen removed by physical stripping is much limited. A simple experiment using nitrogenwithinstead of hydrogen for theofmem-the brane reactor a high packing density
~
•0_ ~.0t-~| a.0~-5 ~
'
QL=l.67xI0-6
3551
.
i 0.3
.
,
.
.
.
o.e
Surface area r a t i o ,
,
.
i
o.g
,
la
As/Am
Fig, 5. Plot for determining mass transfer coefficients, K B t " and K s.
Table 2. Regression results of Fig. 5 Water flow QL (cmS/min)
Ke1~× 106 (m/s)
*Ks × 10~ (m/st
R2 (%)
100 150 200 250 300
2.59 3,38 4.37 5.30 6.25
3.41 4.07 4.27 4,5 l 4.73
98.3 97.3 97.2 96.6 96.1
*Ks values are calculated at q = 0.7.
K. LI et al.
3552
Table 3. Shell-side mass transfer correlations Reference Yang and Cussler (1986) Prasad and Sirkar (1988, Ahmed and Semmens (1992) Costello et al. (1993)
Lietal.(1994)
Range of Reynolds number
Correlations
(
Sh = 1.25 Re
Sh = 5.8(1- ~p)(~)Re°6Sc
Sh = 0.0104 Re°'s°6 Sh
0.1 < Re
SC TM
< 1000
0 < Re < 500
TM
500 < Re < 15,000
S C 0'33
= (0.53~0.581~)Re°'53Sc T M
-
/ d ",~o.967 Sh= 1.164~Rez) Sc °'33
-
10 < Re < I000
de = (4 × flow area)/(total fibre circumference).
catalyst indicates that the dissolved oxygen removal is much reduced. This is understandable as the high packing density of a catalyst tends to reduce the effective area of membrane exposed to water, and hence reduces the dissolved oxygen removed by means of physical stripping. As mentioned in the preceding paragraph, when the area ratio, As~Am, is high, contribution of dissolved oxygen removal by physical stripping becomes unimportant, and the dissolved oxygen removal is mainly achieved by the chemical reaction. Figure 6 shows the effect of water flow rate on the mass transfer coefficient for catalytic reaction using the area factor, ~/, as a parameter. The values of Ks are plotted vs water flow rate using log-log co-ordinates. The hydrogen flow rate in the fibre lumen was maintained at a constant rate of 401/h. As would be expected, for a given water flow rate, the mass transfer coefficient, Ks, is decreased with an increase of area factor, ~/, The regression analysis shows that the best fitted lines drawn through the data points correspond to the variation of Ks as 0.28 to 0.31 power of water flow rate. The increased mass transfer coefficient with water flow rate suggested that the removal of the dissolved oxygen due to the catalytic reaction in the membrane reactor is primarily controlled by the resistance at the liquid-solid boundary layer, i.e. I/ks. This conclusion drawn from the above experimental results is consistent with that of earlier observations (Tay, 1994) for the catalyst bed (randomly packed) reactor used for removal of the dissolved oxygen, The membrane reactors employed in the present study may be viewed as an expanded bed reactor where two different motion behaviours of the catalyst particles were observed in the reactor depending on the a m o u n t of the catalyst used and the water flow rate. These two behaviours are illustrated in Fig. 7. As can be seen from the figure, at a low volume of packed catalyst and a high water flow rate, the catalyst particles are disturbed so that expanded or jiggling bed behaviour is observed [Fig. 7(A)]. When the volume of the packed catalyst particles is high and water flow rate is low, the particles are not disturbed; in this case,
sOoox~o - s Module 1
10o0 = "~ b~ ~
_:
~. ' ° ' ~ ----L--~ ---
100
'r/-0.4
-
~
~
~-1.o
10
u u
~ u,du~,z ~ 10oo ~ .-.------.---loo ~
lo
,~-o.1
-
_
2
3
_
~-0.4 ~-'~i~
5
7.
10x10
-8
Woter flow rote, m 3 / s Fig. 6. Effect ofwater flow rate on the observed mass transfer coefficient, K s.
the membrane reactor acts like a fixed bed [Fig. 7(B)]. The value of the mass transfer coefficients in fluidized bed or fixed bed reactor is often expressed as dimensionless correlations. The appropriate correlations give the Sherwood number as a function of the Reynolds n u m b e r and Schmidt n u m b e r as shown:
Sh = A Re=Sc ~
(13)
or (~-~Z)
(vLdpy(v'~ = A \ - ~ - - / \~-~/
(14)
where A, ¢ and fl are adjustable parameters. Since, in this study, the Schmidt number, Sc, was not varied, a 0.33 power dependence, which is commonly accepted in the literature, was assumed. The values of A and ¢ were obtained by least-square analysis. The
Removal of dissolved oxygen in ultrapure water production
high fluid
~
3553
.,~ep' low fluid
expanded bed behaviour
A
13 Fig. 7. Catalyst particle motion in the membrane reactor.
normalized experimental data have been presented in different ways shown in Figs 8,9 and 10, and is discussed below. Figure 8 plots the mass transfer coefficient, Ks, as Sherwood number, Sh, vs Reynolds number, Re, in log-log form, while Fig. 9 shows a plot which takes account of the area factor, q, and membrane packing fraction, ~b, in the reactor. It can be seen from Fig. 8 that the reactors with either uniformly packed or randomly packed hollow fibres have a similar performance indicating that the way in which the hollow fibres are packed in the reactor is not important as long as the catalyst packed on the shell side surrounds the hollow fibres in the reactor. The general trends of the lines in Fig. 8 are similar to those correlations available in the literature for both fixed and fluidized bed reactors (Chu et al., 1953; Frantz, 1962; Upadhyay and Tripathi, 1975; Dwivedi and Upadhyay, 1977}. If the modified Sherwood, Sh' = {ksdp/DB)fl, and Reynolds, Re' = (vLdp/v) (1 4'), numbers are used, a more general dimensionless correlation can be obtained, and is shown in Fig. 9 where the regression results have been shown to be best described by -
-
Sh =
5.05
[Re(l -
~)]0.307
(15)
50o m ~
• Modore~.~=0.~ • Mod~2, ~=o.~
%~ o o ~ ~.
c II
~ ~g o ~
0
~
,
• Module2, ~=~.o
~
~o
20
~o
Reynolds number, (VLdp/V)
Fig. 8. The observed variation in Sherwood number, Sh, with Reynolds number, Re, for dissolved oxygen removal in the membrane reactors. -% .~. ~
/
,'F 1o i
~
Mod~,e~
]
Module2
" ~'-"-f-
-g. ~= ~ ~
t [
sh'~=5"os(Re) °'3°7
to~
/
R2=97'8~
~g
or
Modified Reynolds number, Re=(VLdp/V)(1-*)
Sh = 0.40 [Re(l -- (~1)]0"307 Sc O'33
(16) r/~ Equation (16) treats the void fraction as a constant value (under fixed bed condition), which may not be the case if the reactor with low catalyst packing densities is operated with a high water flow rate. As mentioned earlier, under certain operating conditions, the catalyst bed is expanded and this will inevitably increase the void fraction of the catalyst bed. The following linear relation may be used to correlate the catalyst bed expansion: Z 1 ~:i = el~-/,
~:i < 1
t17)
Fig. 9. The observed variation in modified Sherwood number, Sh', with modified Reynolds number, Re', for dissolved oxygen removal in the membrane reactors.
where e~ and Z~ are void fraction and length of the catalyst bed under fixed bed conditions, respectively, while e~ and Z~ are void fraction and length of the catalyst bed under expanded bed conditions, respectively. By incorporation ofe~ into the Reynolds humber, the modified Reynolds number, Re" is defined as Re" =
l!Ldp(1 -- q)) v(l - - el)
(18)
K. Ll et al.
3554
Plotting the modified Sherwood number, (ksdp/OB)q, vs the modified Reynolds number, Re", shown in Fig. 10, the regression results give 0.55 ~ R e ( 1 - q~)]°22 Sh = SC O'33 (19) L 1--- ~i _l '
Figure 11 shows the effect of water flow rate on the dissolved oxygen concentration in the product stream. The experimental data were obtained using reactors 3 and 4 connected in parallel. The total volumes of the catalyst packed in these two reactors are 400 cm3; thus, the area ratio, As/A,, is very high
The correlations in eqs(16) and (19) have similar trends to many other available correlations suggested earlier as shown in Table 4. The mass transfer coefficient Ks varies as a 0.22 power with Re", which is consistent with the 0.22 power (Chu et al,, 1953) and the 0.21 power (Upadhyay and Tripathi, 1975) found by previous workers. In comparison with Sh vs Re', the mass transfer coefficient, Ks, is found to vary with the 0.307 power with Re', which is slightly different from that of 0.28 given by Dwivedi and Upadhyay (1977). The numerical constants found here cannot be directly compared with those given in Table 4 as, in this study, the area factor, q, is an unknown value which dictates the final numerical constants. If the constants listed in Table 4 are assumed to be true, the area factor of the catalyst used in this study is around 10-40%.
and is equal to 10. Under such a condition, K B ~ K s , and eqs (1) and (15) can therefore be used for prediction purposes. The calculated data are shown by the solid curve. As can be seen from Fig. 11, the observed experimental data and the calculated results are in good agreement. The dissolved oxygen level in the outlet stream can be achieved as low as 1.2 ppb, which is basically impossible to achieve using conventional physical methods. CONCLUSIONS A novel membrane reactor has been developed for removal of the dissolved oxygen in water from parts per million (ppm) to a level less than 1.5 ppb. The membrane reactor comprises a microporous hollow fibre module packed with a palladium catalyst in the void space. Mass transfer studies show that the over-
~ o_
:~u~ ~* -~ "
• Module I • Module 2
lo
• Exp. data -Calculated values
•
~
R2=96.1% 1
~0
~ ~~ ~_ 50 ~ E 40
/
~ .E
10
~'~
0
4
Modified Reynolds number, R'e=(VLdp/#)(1-$)/(1-~i)
8 Water
Fig. 10. The observed variation in modified Sherwood num-
' 16
12
20x10
-6
mS/s
flow fete,
Fig. 11. Comparison between experimental results obtained
bar, Sh', with modified Reynolds number, Re", for dissolved fromreactors 3 and 4 and calculated results from eqs (1) and oxygen removal in the membrane reactors.
(15).
Table 4. Comparison of correlations Reference Chu et al. (1953) Frantz (1962)
Correlations
/ Re ~T M Sh = 5. ~l_e.i -Sc T M [ Re \ °'3°
Sh = 5 \11--1ei.}
Sc°33
Range of Reynolds number
1
< ( Re ~ \I - eiJ < 30
1
< ( Re ~ \1 - eJ < 80
(Re~
(Re) ~
Upadhyay and Tripathi (1975)
Sh = 3.8155k1-~/) T M Sc°'Sa
Dwivedi and Upadhyay (1977)
Sh = 1.1068 ReO.28SCO.33
Re < 10
This study
Sh = 0.40 ERe(1 -- q~)]o.3o7Sco.aa
Re(1 - ~b)< 10
= n k ~ J
\
l-e,
< 20
)
50
Removal of dissolved oxygen in ultrapure water production all mass transfer coefficient for physical stripping, KsL, is generally smaller than that for chemical reaction, Ks. Both KBL and Ks are dependent on the water flow rate indicating that the liquid films adjacent to the hollow fibre membrane and the solid catalyst surface control the removal of the dissolved oxygen. The mass transfer study further reveals that at high catalyst loadings, the total oxygen removal is mainly achieved by the chemical reaction alone and the removal of the dissolved oxygen by physical stripping is negligible. Under such a condition, the hollow fibre membrane employed becomes akin to a gas distributor and ensures a constant supply of purified hydrogen for the catalytic reaction. Thus the membranes provide a stationary interface which enables both the gas and liquid flows to be operated independently and operation constraints such as flooding and loading points, usually found in packed columns, no longer exist. Based on the experimental results, a correlation has been developed, which can be used for the design of membrane reactors of the type described in this paper. Under high catalyst loadings, the correlation proposed can be used alone for design purposes, whilst at low catalyst loadings, i.e. both chemical reaction and physical stripping are significant for the dissolved oxygen removal, the correlations listed in Table 3 and the correlation developed must be employed together for the design of the membrane reactor.
Acknowledgements--The authors are grateful to the Singapore Science Council for the award of RDAS Grant (ST/90/03) to support the research and development work on ultrapure water production. NOTATION A At A,~ As C8
constant in eqs (13) and (14) effective catalyst surface area, m 2 total membrane area, m 2 total surface area of catalyst, m 2 dissolved oxygen concentration, kmol/m 3
do dp HB k
outside diameter of hollow fibres, m catalyst particle diameter, m Henry's law constant individual mass transfer coefficient
kns KB
reaction rate constant, m/s overall mass transfer coefficient of oxygen defined in eq. (12), m/s overall mass transfer coefficient of oxygen for physical stripping, m/s
KnL Ks n NB
overall mass transfer coefficient of oxygen for catalytic reaction, m/s number of fibres rate of the dissolved oxygen transfer, kmol/s
Q Re Re' Re"
fluid flow rate, m3/s Reynolds number ( = vcdp/v) modified Reynolds number(= vcd~/v)(1 - ~b) modified Reynolds number ( = vLdp/v) (1 -q~)/(1 -- el)
Sh Sh'
Sherwood number ( = Ksdp/DB) modified Sherwood number (= Ksdp/DB)q
CE$ 50-27-E
vL Z Z1 Zi
3555
velocity of water in membrane reactor without catalyst packings, m/s effective length of the fibres, m length of catalyst bed at fixed bed condition, m length of catalyst bed at fluidized bed condition, m
Greek letters ct fl ¢ el
parameter in eqs (13) and (14) parameter in eqs (13) and (14) porosity of catalyst bed void fraction of fixed catalyst bed
e.i q v q~
void fraction of fluidized catalyst bed area factor, defined in eq. (6) kinematic viscosity, m2/s packing fraction of hollow fibres
Subscripts 1 physical stripping 2 catalytic reaction g gas i interface L liquid m membrane S catalyst surface Superscripts in reactor inlet out reactor outlet . liquid phase in equilibrium with gas phase
REFERENCES Ahmed, T. and Semmens, M. J., 1992, Use of sealed end hollow fibres for bubbleless membrane aeration: experimental studies. J. Membrane Sci. 69, 1-10. Bayer AG, 1990, Catalytic removal of dissolved oxygen from water. Organic Chemical Division, Forschung/Anwendungstechnik Lewatit, Ionenaustauscher/Katalysatoren, Geb. B106, Leverkusen, pp. 1~. Chu, J. C., Kalil, J. and Wetteroth, W. A., 1953, Mass transfer in a fluidized bed. Chem. En#ng Prog. 49, 141-149. Chua, I., 1994, Elimination of dissolved oxygen in ultrapure water by chemical reduction and physical permeation processes. M.Eng thesis, Department of Chemical Engineering, National University of Singapore. Costello,M. J., Fane, A. G., Hogan, P. A. and Schofield, R.W., 1993, The effect of shell side hydrodynamics on the performanceof hollow fibre modules. J. Membrane Sci. 80, 1-11. Dwivedi, P. N. and Upadhyay, S. N., 1977, Particle-fluid mass transfer in fixed and fluidizedbeds. Ind. Engn# Chem. Des. Dev. 16, 157-165. Frantz, J., 1962, Design for fluidization. Chem. Engn9 69, 161-178. Imaoka, T., Yagi, Y., Kasama, Y., Sugiyama, 1., Isagawa, T. and Ohmi, T., 1991,Advanced ultrapure water systemsfor ULSI processing, in Proceedinos of the lOth Annual SemiconductorPure Water Conference, Santa Clara, CA, 26-28 February (Edited by M. K. Balazs), pp. 128-146. Kasama, Y., Yagi, Y., Imaoka, T., Kawakami, M. and Ohmi, T., 1990, Advanced D.I. water system with low dissolved oxygen for ULSI processing. Proceedinos Institute of EnvironmentalScience, pp. 344-349.
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Suppiah, S., Kutchcoskie, K. J., Balakrishnan, P. V. and Chuang, K. T., 1988, Dissolved oxygen removal by combination with hydrogen using wetproofed catalysts. Can. J. Chem. Engng 66, 849-857. Tai, M. S. L., Chua, I., Li, K., Ng, W. J. and Teo, W. K., 1994, Removal of dissolved oxygen in ultrapure water production using microporous membrane modules. J. Membrane Sci. 87,99-105. Tay, Y. T., 1994, Removal of dissolved oxygen using random packed catalyst bed. Final year, B.Engng Research Project, Department of Chemical Engineering, National University of Singapore. Upadhyay, S. N. and Tripathi, G., 1975, Liquid phase mass transfer in fixed and fluidized beds of large particles. J. Chem. Engno Data 20, 20-26. Yang, M. and Cussler, E. L., 1986, Designing hollow-fibre contactors. A.I.Ch.E.J. 32, 1910-1916.