Oxidation of Ce(III) to Ce(IV) by Ozone in Nitric Acid Medium Using a

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Oxidation of Ce(III) to Ce(IV) by ozone in nitric acid medium using a static mixer: Mathematical modeling and experimental validation Sukhdeep Singh, Tessy Vincent, Ashok N Bhaskarwar, and Trushit Makwana Ind. Eng. Chem. Res., Just Accepted Manuscript • DOI: 10.1021/acs.iecr.7b03163 • Publication Date (Web): 07 Sep 2017 Downloaded from http://pubs.acs.org on September 12, 2017

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Industrial & Engineering Chemistry Research

Oxidation of Ce(III) to Ce(IV) by ozone in nitric acid medium using a static mixer: Mathematical modeling and experimental validation Sukhdeep Singh1*, Tessy Vincent1, Ashok N. Bhaskarwar2, Trushit Makwana1 1

Process Development Division, Nuclear Recycle Group, Bhabha Atomic Research Centre, Trombay, Mumbai, 400085, India 2

Department of Chemical Engineering, Indian Institute of Technology, Hauz Khas, New Delhi, 110016, India

Abstract. This paper deals with a cerium based redox chemical process which is employed to remove fixed radioactive contaminants from alpha bearing metallic surfaces, worldwide. Oxidation of cerium to its fourth valence state is a crucial step of this process. We present a bench scale study to obtain the effect of various process parameters on conversion of Ce(III) to Ce(IV) by ozone in nitric acid medium, using a SMX type of static mixer. Two different mathematical models, one considering a simple semi-batch mixed reactor configuration, and another based on the concept of axial dispersion were developed for predicting Ce(III) conversions using the said static mixer, in a full recycle mode. Simulation results from the models matched fairly well to experimental data, with respective mean absolute percentage deviation of 21.1% and 17.4%, computed over all the parameters. An attempt to extend these models for other non-nuclear environmental applications is also described. Keywords: Oxidation of cerium; Ozone; Static mixer; Mathematical modeling; Nuclear decontamination.

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1

Introduction Stainless steel, owing to its excellent corrosion resistance and radiation damage bearing

capacity, is one of the most widely used construction materials for nuclear installations. During the operation/decommissioning of reprocessing plants and fuel fabrication facilities, significant amount of alpha-contaminated metallic waste (alpha activity > 4000 Bq/g) is generated. Research efforts are being put to establish adequate decontamination techniques for safe handling of such waste. Cerium redox decontamination process (Figure 1) is one such technique which involves the contacting of contaminated metallic components with a solution of Ce(IV) in HNO3. Ce(IV) oxidizes the metal surface ( E 0 = 1.61 V versus NHE), thereby releasing the fixed radio nucleotides and steel constituents (Fe, Cr and Ni) into the solution, and itself gets reduced to Ce(III)1,2. Redox reactions of Ce(IV) occurring on the surface of contaminated steel components are reported elsewhere2. Treated components, after a typical removal of 10-15 µm of the top layer, could then be placed into non-alpha waste category2. The reduced cerium i.e. Ce(III) so generated is then oxidized back to Ce(IV) by reaction with ozone ( E 0 = 2.07 V versus NHE), so as to maintain the required Ce(IV) concentration, hence sustained corrosion rates and a lower treatment time. This oxidation of reduced cerium occurs as per the following chemical reaction3.

O3 + 2Ce( III ) + 2 H + → 2Ce( IV ) + H 2 O + O2 Therein, dissolved ozone also undergoes a self-decomposition reaction to molecular oxygen. Kinetic rate equations of these two chemical reactions have been reported by Singh et al.4 and Mizuno et al.5 respectively.


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Figure 1. Pictorial representation of cerium redox decontamination process, showing the dissolution of radio nucleotides and base metal, along with simultaneous regeneration of Ce(IV) by ozone. At this juncture, a question may arise, as to why one cannot decontaminate the metal directly with ozone. The answer to this lies in the fact that, firstly, the direct oxidation of stainless steel by gaseous ozone is a very slow process. Pohjanne and Sitala6 have reported the corrosion behaviour of AISI 316L and AISI 304L steels in the ozone environment. As per there studies, under the conditions: T = 20-30 °C, P = 1000 kPa, O3 = 9.5% and gas velocity of 5 m/s (over the steel components), corrosion rates of about 0.017 mm/year (0.00194 µm/hour) and 0.012 mm/year (0.00136 µm/hour) were observed for 304L and 316L steels, respectively. These rates are about 1000 times lower in comparison to those observed in Ce(IV) mediated corrosion (2-3 µm/hour)1 of the present process. Secondly, oxidizing the top layer of alpha-contaminated metallic components directly with ozone will generate air-borne activity, which is not welcome in nuclear facilities. Lastly, Ce and dissolved radionuclides, i.e. Pu and Am can be easily 3 ACS Paragon Plus Environment


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recovered from the spent solution generated in cerium redox process by using standard extractants2. Such a recovery of radionuclides from gas phase would not be a feasible option, if direct ozone decontamination is employed. Ceric regeneration forms an important part of this process. When effected in nitric acid media, a preferred choice in the nuclear fuel cycle owing to established vitrification practices, it has a crucial role to play in the design and scale up of decontamination facilities. In such assignments, one has to estimate the rate (kmole/s) at which regeneration of Ce(IV) in the decontamination solution is required, depending upon the area of contaminated metallic components charged to the facility and the process conditions. Regeneration rate of Ce(IV), so obtained, further decides the capacity of ozone generation unit required for the facility. Besides this in-situ conversion of Ce(III) to Ce(IV), the oxidation of cerium present in the secondary waste is again required for its selective recovery as Ce(IV) (for reuse) by tri-butyl phosphate from among the radio nucleotides (Pu and Am) and the steel corrosion products (Fe, Cr and Ni), towards the culmination of process2, as mentioned earlier. In addition to the targeted decontamination related applications, there exist many other uses of Ce(III)/Ce(IV) redox couple also, wherein this regeneration of Ce(IV) may be required. An excellent general review of those applications is provided by Arenas et al.7 and Binnemans8. Studies specific to destruction of various organic pollutants e.g. phenols, ketones, carboxylic acids, diazols and alkyl sulfonic acids, using Ce(IV) as an oxidant, are well reported in literature923

. Etching of chromium, using Ce(IV) as an oxidizing agent, has been employed in thin film

transistor (TFT)-liquid crystal display (LCD) manufacturing industry; and spent Cr-etching solutions have been treated electrochemically for regeneration of Ce(IV)24,25. Besides this,


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gaseous pollutants e.g. SO2/NOx26 and H2S27 could also be removed from the waste streams, using Ce(IV) assisted processes. Valency change of cerium could be effected either electrochemically or chemically with ozone. Electrochemical regeneration results into the generation of hydrogen at cathode28,29, posing a safety concern, particularly in nuclear decontamination. In addition to hydrogen, nitrogen oxides (NOx) also get generated during the process involving a nitric acid medium30, which requires a separate scrubbing system for regulated release to environment. Due to these issues, and an easy generation of ozone from ambient air, ozone based route for regeneration of Ce(IV) is usually adopted. This ozone based decontamination process has received the attention of several researchers. Ponnet et al.1 conducted the decontamination of a steam generator of a pressurized heavy water reactor (PHWR) using Ce(IV) in sulfuric acid media, and regeneration with ozone. The equipment was successfully decontaminated up to the desired release levels of radioactivity. Carie et al.3,31 pioneered the decontamination process for AISI 304 L steel by Ce(IV) in 4 M nitric acid (instead of sulfuric acid), in which the reduced cerium was also re-oxidized by ozone. In the first part of their work3, authors reported a decontamination study targeted at obtaining a corrosion rate of ≈ 10 µm per day. In their second paper31, process parameters were optimized for the decontamination of AISI 304 L steel. Empirical equations for estimating corrosion rates, corroded mass, metal solution potential difference and corrosion current density were reported. Recently, Shah et al.2 conducted the decontamination of alpha-contaminated SS304L samples, using cerium in nitric acid media, at laboratory scale. As per their study, Ce(IV) assisted with ozone based regeneration was found to be effective for decontamination of active stainless steel coupons. Extraction experiments using conventional plutonium uranium redox extraction (PUREX) solvent



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i.e. 30% tributyl phosphate (TBP) in n-dodecane, for the recovery of Pu(IV) and Ce(IV) from the spent solution, were also performed by the authors. Am(III) was recovered by employing 0.2 M tetra (2-ethylhexyl diglycolamide), commonly referred as TEHDGA, in 30% isodecyl alcohol. All these studies, reported so far on this topic, have definitely provided an insight into the process; but none of them has separately addressed the Ce(IV)-regeneration part of this process, mathematically and experimentally. Due to these reasons, the need for present study was felt. In this work, the ozone based oxidation of Ce(III) to Ce(IV) in nitric acid medium has been studied systematically, on a bench scale facility using a polyvinylidene fluoride (PVDF) made static mixer with open blade structure (equivalent to Sulzer SMX) as a gas-liquid contactor. Effect of each of the process parameters viz. ozone gas concentration, Ce(III) initial concentration, nitric acid concentration, temperature, liquid and gas flow rate is discussed. Furthermore, two mathematical models, one based on simple semi-batch reactor configuration and the other using concept of axial dispersion, have also been developed and duly validated against the experimental data. 2 2.1

Experimental Chemical reagents Cerium (III) nitrate hexahydrate (analytical grade, supplied by M/s. Orion Chemicals,

Mumbai) was used as the source for Ce(III) ions. Plant grade nitric acid (12 M) was used for preparing the cerium solutions in demineralized water. Other chemicals e.g. ferrous ammonium sulphate (analytical grade) and ferroin indicator solution (1% of 1,10 phenolphthalein in ethanol, supplied by Merck Life Sciences Pvt. Ltd., Mumbai) were used for titrimetric analysis of aliquots.


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2.2

Experimental setup The experimental set-up used in the study is shown in Figure 2.

Figure 2. Schematic of bench scale experimental setup. Ambient air was first concentrated in an oxygen concentrator to give about 95% O2, employing pressure swing adsorption. Oxygen stream so obtained was then fed to a corona dischage based water-cooled ozone generator (Make: Chemtronics, India, Capacity: 50 g/h of O3) to produce a mixture of O3 and O2. A tapping, with gas flow rate of ≈ 0.1 L/min, from main ozone line was taken to an ozone anlayzer (Make: BMT Messtechnik GmbH, Germany, Model: BMT-964) through a needle valve, to get the ozone concentration of the incoming feed gas, in



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g/m3 under normal conditions of temperature and pressure (0 °C and 1 atm, abbrebated as NTP here onwards), whenever required. Main gas stream from ozonator was then injected into the liquid stream, containing a solution of cerium(III) nitrate hexahydrate in nitric acid, using an ozone injector. The injector was made to suck the gas by the vacuum created due to circulating liquid (at 11 L/min in all experiements, except in those for studying the effect of liquid flow rate), because of Bernoulli effect. Oxidation of Ce(III) to Ce(IV) was happening in the following SMX static mixer (D = 1.9×10-2 m and L = 0 .2 m), upon intimate mixing with ozone. Such a contactor, with high specific interfacial area and no moving parts, was chosen with the aim of better performance and minimum maintenance during the intended deployment of process in future. Figure 3 shows the typical geometry of an internal mixing element of a Sulzer SMX static mixer.

Figure 3. Sample geometry of the mixing element of a Sulzer SMX static mixer (Reprinted with permission from Hirschberg et al.32. Copyright 2009 Elsevier.). The gas-liquid stream coming out of the static mixer was fed to a cylone separator to give a liquid stream having regenerated Ce(IV) and another gas stream of unreacted ozone. Unreacted ozone was directed to a catalytic destructor, having a bed of CuO and MnO2, for converting it


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back to molecular oxygen. Gas, after the destruction of ozone, was vented to atmosphere through air ejector-I. This ejector served the purposes of providing a dilution, and a driving force, to the off-gas, thereby taking O3 concentration beyond the thrushold limit value of 0.1 PPM, and of compensating for the pressure drop in catalytic bed of ozone destructor. Concentration of ozone in the off-gas coming out of the destructor was monitored contineously by ozone anlayzer. The analyzer was provided with a second air ejector, to draw the low pressure off-gas from the destructor. Air flow rate to both the ejectors was so adjusted, that there is sufficient flow rate ( ≈ 0.1 L/min) of off-gas from the destructor to the ozone anlayzer for accurate and live monitoring. On the other hand, Ce(IV) rich stream from gas-liquid separator was fed into a solution tank of 5 L capacity. This tank was equipped with a glass-walled oil bath (non-flammable transformer oil), having an electric immersion element of 1.5 kW, for heating the solution. Another oil-thermowell made of glass was used for hosting a resistance temperature detector (RTD) for monitoring the temperature of solution. Feedback signal from RTD was supplied to a controller for regulating the set temperature. A vent line, connected to air ejector-II through a needle valve, was also provided to this tank. This line was meant for avoiding any pressure buildup in the solution tank due to a smaller fraction of gas bubbles escaping from the gas-liquid separator. Solution from this tank was again recycled back to the loop via. an air operated double diaphragm (AODD) pump, for further conversion. This way, the liquid in the loop happened to be in full recycle (batch), whereas the gas, in a contineous mode; thus making the overall process a semi-batch operation. Material of construction for piping (20 NB, schedule 80) in the loop was polyvinyl chloride (PVC). Polymers material were choosen to minimize the corrosion attack by Ce(IV) in HNO3.



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2.3

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Analytical methods Volumetric titrations with freshly prepared 0.01 M ferrous ammonium sulphate, using

ferroin as redox indicator, were employed for estimating Ce(IV) concentrations in the aliquots drawn at regular intervals. Samples were analyzed at the end of experiments, in order to allow the decay of µM-level dissolved ozone into oxygen, thereby eliminating any possibility of the consumption of titrant by ozone. Ce(III) (the reactant) concentrations were then obtained by subtracting the Ce(IV) concentration from the initial Ce(III) concentrations. Ce(III) conversions at various times were subsequently computed from the standard definition of conversion as per Eq. (1).

 [Ce( III )](t )   X (t ) = 1 −  [Ce( III )]initial  3

(1)

Mathematical modeling

As described earlier, two mathematical models have been developed in this study for batch mode regeneration of Ce(IV) by ozone using a SMX static mixer. The following sections present the formulation. 3.1

Assumptions Following simplifying assumptions were made while modeling the experimental loop

under study, and numerical simulations of the same: a) Liquid phase was assumed to be in full recycle or batch mode, and the gas being in continuous mode.


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b) Same gas holdup ( ε G ) values, those predicted by the available correlations for Sulzer SMX static mixer in once through mode, were adapted for modeling the batch mode operation of liquid in the present study. c) Gas phase mass balance for ozone in both the models was obviated, owing to a small length of static mixer (= 0.2 m) and a full recycle operation, wherein fresh ozone was being fed to the mixer at every recirculation cycle, and therefore only a negligible change in gas concentration was expected. Unavailability of gas phase Péclet number data for SMX static mixer from the literature was another reason for omitting this gas phase mass balance. However, a satisfactory match with experimental data justified the assumption, as anticipated. d) For having a simple numerical solution of the models, the value of mass transfer enhancement factor E , expressions for which involve complex implicit equations in terms of Hatta number and enhancement factor for infinitely fast reaction E∞ , was taken to be unity; however, it was included in the model equations for the sake of completeness. This assumption could be justified because the maximum pre-calculated value of E , under one of the best experimental conditions ([Ce(III)] = 0.1 M, [HNO3] = 4 M, G0

= 29 ± 1.5 g/m3 (NTP), QG = 5 L/min, QL = 11 L/min and TC = 35.7 °C), was only 1.0504 (calculation shown in Supporting Information).

This value was estimated at the

beginning of reaction, under the aforementioned process conditions, and it would approach unity as the reaction proceeded. e) Marginally decreasing nitric acid concentration during reaction, has a negligible effect on the self-decomposition of ozone. Ozone self-decomposition rates were therefore 11 ACS Paragon Plus Environment


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evaluated only at the initial pH values, in the semi-batch model simulations. However, in the axial dispersion model, the values of nitric acid concentration, hence the pH , were updated after each cycle of solution algorithm for a more rigorous solution. f) No heat effects due to chemical reactions were accounted for in the models. This assumption could also be justified, as noticeable temperature changes were not observed during the experimental runs. 3.2

Semi-batch model (SBM) With the assumptions described in Section 3.1, schematic of an equivalent system

modeled is shown in Figure 4, wherein VC is the volume of contactor, which is the sum total of liquid volume V L and the volume of entrapped gas VG .

Figure 4. Schematic of a semi-batch system equivalent, modeled mathematically.


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Governing differential equations for O3, Ce(III), and HNO3 (designated as A, B and C for simplicity) were written from the general mass balance which, for dissolved ozone in the liquid phase, could be written as: O3 = O 3 − O 3 − - O 3 O 2 Mathematically, dn A = (N A dt

z =0

) A − rAV L − rd V L

(2)

, where the expressions for each of the unknown variables are described below, starting from the left side of equation. Moles of ozone n A , and flux of ozone at interface N A

z =0

, are given by Eqns. (3-4).

nA = VLC A

NA

z =0

= k L E (C AS − C A )

(3) (4)

Solubility of ozone in water as a function of temperature and pH is described by Eqns. (56)33.

 − 2428  H = 3.84 × 10 7 (10 pH −14 ) 0.035 exp   T  C AS = 55.56

pA H


(5)

(6)


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Value of partial pressure of ozone p A (atm) in Eq. (6) was obtained (Eq. (7)) from the ozone concentration indicated by analyzer (in g/m3 at NTP) after due corrections (please refer Supporting Information) for actual temperature and pressure in the experiments, as per ideal gas equation. p A = 4 .6693 × 10 − 4 G 0

(7)

Mass transfer enhancement factor E was assumed to be 1 as per assumption (d). Next unknown of Eq. (2), the total interfacial area A, is expressed as the product of specific interfacial area

a

and contactor volume VC (Eq. (8)), A = aVC

(8)

V  VC =  L  εL 

(9)

where VC is given by Eq. (9).

Then moving to the reaction terms of Eq. (2), the rate of consumption of dissolved ozone by Ce(III), as reported in our earlier kinetics study4, is provided by Eq. (10). (10)

− r A = kC An C Bm C Cp

, where the values of n, m and p are 1.21, 0.68 and 0.54 respectively. Rate constant k in Arrhenius form is expressed as per Eq. (11).

k = 6.29 × 10 6 exp(

− 5231.4 ) T

(11)

The second order self-decomposition rate equation for ozone, as reported in Mizuno et al.5, is given by Eq. (12).


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[

− rd = 48 × 10 3 × 37.4 10 ( pH −14)

]

0.73

exp(

− 13367(T − 293) 2 )C A 293T

(12)

After all these substitutions, the following equation of ozone mole balance in liquid is arrived at

 V d (VL C A ) = k L (C AS − C A )E  a L dt  εL

  − kC 1A.21C B0.68 CC0.54VL − k d C A2 CC0.73VL 

(13)

On dividing it throughout by VL , we have

 1 dC A = k L a(C AS − C A )E dt εL

  − kC1A.21C B0.68CC0.54 − k d C A2 CC0.73 

(14)

Expressions for volumetric mass transfer coefficient k L a and liquid holdup ε L , for a Sulzer SMX type of static mixer in Eq. (14), were adapted from an air-water hydrodynamics study reported in Mouli et al.34; and are given by Eqns. (15-17). Eq. (15) was used after having dual corrections (Eq. (18)) for ozone-water system by multiplying it with ratio of diffusivities as per film theory of mass transfer35. (k L a ) Air −Water = 0.003(

ε G = 2.12 × 10 −5 (

Du L ρ L

µL

Du L ρ L

µL

) −0.06 (

) 0.099 (

Du G ρ G

µG

Du G ρ G

µG

) 0.0307

)1.18 exp(0.01N e )

ε L = (1 − ε G ) kLa =

 D D A  DO2 ×  (k L a ) Air −Water  = A (k L a ) Air −Water DO2  D Air  D Air

(15)

(16)

(17)

(18)

In these equations, density of gas ρG was substituted by the density of actual ozoneoxygen mixture (maximum 4% O3, rest 96% O2 on wt./wt. basis), which was estimated from that 15 ACS Paragon Plus Environment


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of pure oxygen as per ideal gas equation (Eq. (19)). Similarly, for obtaining the viscosity µG of this gas mixture, Sutherland’s formula36 was used for pure oxygen as per Eq. (20),

ρG =

PM O3

(19)

RT  T 3/2   T + C 

µ G = λ × 10 − 6 

(20)

, where λ and C being the constants specific to oxygen, with respective values of 127 K and 1.693411300 Pa.s/K1/2. For the liquid phase, the value of density ρ (in g/cm3) and viscosity µ (in cP) were calculated from the regression equations37 (Eqns. (21 and 22)), fitted into data38,39 for nitric acid solutions in water as a function of nitric acid mass fraction

x

and temperature T , useful in the

range 0 < x < 0 .4 and 22 < T < 70 °C. Nitric acid mass fraction x was substituted in terms of concentration and temperature as per Eq. (23) in both the equations for density and viscosity. The implicit form of Eq. (21) was then solved numerically to get ρ . The value of ρ so obtained, was then fed into Eq. (22) to obtain µ . Those values of ρ and µ

were further

converted to metric units as per Eqns. (24 and 25), and were subsequently used in Eqns. (15 and 16) for (k L a ) Air −Water and ε G estimation.



ρ = (1 + 0.6741x1.0988 )exp49.20 − 



1 T

1   293 

µ = (1 + 3.2043x1.9037 )exp1723.24 − 

x=

1 T

1   293 

CC M C 1000ρ 16

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(21)

(22)

(23)

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ρ L = 1000 ρ

(24)

µ L = 1 × 10 −3 µ

(25)

Effect of addition of cerium salt on density and viscosity of nitric acid solution was neglected, as the Ce(III) concentrations employed were in the range 0.0125 M to 0.1 M only. However, if more rigorous values are sought, experimental measurements can be made40. Diffusivity of air to be used in Eq. (18) was estimated from the Wilke and Chang equation41 (Eq. (26)), taking water as the medium, and that for ozone was computed42 using Eq. (27) below.

7.4 × 10 −8 (σM Air ) T 0 .5

D Air =

1000µ L (V Air )

0 .6

 − 1896  D A = 1.1 × 10 −6 exp   T 

(26)

(27)

On the similar lines as for ozone, the mass balance for Ce(III) and HNO3 in liquid phase was written, by omitting the interfacial mass transfer term from the equation for ozone as shown below: = − Governing differential equations for Ce(III) and HNO3 thus obtained are given by Eqns. (28 and 29). dC B = −2kC 1A.21C B0.68 C C0.54 dt


(28)


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dC C = −2kC 1A.21C B0.68 C C0.54 dt

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(29)

The model constituted by species balance Eqns. (14, 28 and 29) and other auxiliary equations could be solved using following set of initial conditions (Eq. (30)):

t =0

 C A = 0   CB = CB0  CC = CC 0  3.3

(30)

Axial dispersion model (ADM) In this model, for describing the oxidation of cerium by ozone, concept of axial dispersion

was employed. A differential shell balance (Figure 5) was written for a segment of the reactor, and governing differential equations were obtained for the various reacting species.

Figure 5. Schematic of differential shell balance on the static mixer for axial dispersion model (SMX elements not shown).


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At steady state, the various terms of this differential shell balance for ozone could be written as: O3 ℎ + O3 ℎ # + O3 − O3 ℎ − O3 ℎ # − O3 − - O 3 O 2 = 0

Mathematically, we have  dC   dC  − DL (Sε L ) A  + DL (Sε L ) A  − C Al + ∆l (u L S ) + C Al (u L S ) + (S∆l )a[ N A dl  l  dl  l +∆l

z =0

] − (S∆lε L )[rA + rAd ] = 0

(31)

On dividing by S∆ l and taking the limit ∆l → 0 , we got DLε L

(

d 2C A dC A − uL + a NA 2 dl dl

After substituting the expression for ozone flux N A

z =0

z =0

)− ε

L

( rA + rAd ) = 0

(32)

as per Eq. (4), and those for rA and rAd as

per Eq. (10) and Eq. (12) respectively, following dimensionless variables and numbers (Eqns. (33-39)) were introduced.

z=

cA =

l L

(33)

CA C AS

(34)



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ci =

Ci Ci 0

(35)

uL L DL

(36)

St L = k L a τ L

(37)

PeL =

n −1 m D a A = kτ L C AS C B 0 C Cp0

Da Ad

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(38)

K  = 48 × 10 k dτ L C AS  w   CC 0 

0.73

3

(39)

Finally, we arrive at the dimensionless equation (Eq. (40)) for ozone mole balance in liquid as,

ε L d 2cA PeL dz

2

−

dc A + St L E (1 − c A ) − DaA c An c Bm cCp ε L − DaAd c A2 cC0.73ε L = 0 dz

(40)

The correlation for Péclet number in above equation was adapted from the literature34, for a Sulzer-SMX static mixer and air-water system as per Eq. (41). Pe L = 2.9080(

Du L ρ L

µL

) 0.1636 (

Du G ρ G

µG

) 0.0533

(41)

Volumetric mass transfer coefficient k L a , for calculating Stanton number StL , was estimated from Eqns. (15 and 18), as described earlier in the Section 3.2. Liquid residence time τ L in Eq. (37) was computed simply by dividing length of static mixer by superficial liquid

velocity as described by Eq. (42).

τL =

L uL


(42)

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Next unknown of Eq. (40), the liquid holdup ε L , was computed from Eqns. (16 and 17). K w , the ion product constant of water, was used to estimate the

OH

−

concentration (replaced by

pH in semi-batch model) in the solution, and was computed from the regression equation (Eq.

(43)) fitted into K w versus T data43 at atmospheric pressure.

(

K w = 1 × 10 −14 9.699 × 10 −5 (T − 273.15) − 0.005304(T − 273.15) + 0.121(T − 273.15) + 0.1122 3

2

)

(43)

This way, finishing the mass balance for ozone, similar equations for Ce(III) and HNO3 were written in dimensionless form by omitting the interfacial mass transfer term, as

ε L d 2 cB PeL dz

2

ε L d 2 cC PeL dz 2

−

dcB − DaB c An c Bm cCp ε L = 0 dz

(44)

−

dcC − DaC c An c Bm cCp ε L = 0 dz

(45)

, where the dimensionless Damköhler numbers were defined by Eqns. (46 and 47). n Da B = 2kτ L C AS CBm0−1CCp0

(46)

n Da C = 2kτ L C AS CBm0CCp0−1

(47)

This axial dispersion model constituted by Eqns. (40, 44, and 45) and the other auxiliary equations requires two boundary conditions, one at z = 0 and another at z = 1 , for the complete solution. Next section describes these boundary conditions in detail.

3.3.1

Boundary conditions for ADM This mathematical model constitutes a coupled-second order boundary value problem. To

obtain the solution of such a system, one needs two boundary conditions for each of the species



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differential equations, at the physical domain boundaries. Figure 6 depicts the geometry of the reactor, defining the dimensionless domain boundaries of our system corresponding to z = 0 and z = 1.

Figure 6. Schematic diagram of reactor geometry in full recycle (batch) mode, with static elements starting from z = 0 and ending at z = 1. Also shown is the liquid being recycled back after separation from gas at exit. At z = 0 , the boundary condition is that all dimensionless concentrations are known (Eq. (48))

ci

z =0

=1

(48)

, with i = B & C . For A right hand side of Eq. (48) will be zero, as the solution was not presaturated with ozone at the beginning. At z = 1 , we have the boundary conditions44,45 given by Eqns. (49-51).

dc A dz

(

= − f DaA1 c An c Bm c Cp z =1

dcB dz

(

z =1

+ DaA2 c A2 cC0.73

= − f DaB c An c Bm cCp z =1

z =1


)

z =1

)

(49)

(50)

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dcC dz

(

= − f DaC c An cBm cCp z =1

z =1

)

(51)

, with

 1   f = 1 +  PeL 

−10

Pe L → ∞, f → 1( PFR )   Pe L → 0, f → 0( MFR ) 

(52)

(53)

This boundary condition needs a slight explanation. At the end of reactor, the condition is that the gas and liquid are in contact upto z = 1− , and immediately after that, i.e. at z = 1 they are separated, mass transfer ceases, and the liquid gets recycled back. Romanainen and Salmi44 have suggested a boundary condition for gas-liquid tubular reactors in which Péclet number terms in the governing differential equations (Eqns. (40, 44 and 45) in our case) have been dropped, assuming the plug flow conditions at the reactor exit; and then to take the rest of equations as boundary condition corresponding to z = 1 . But this case might not always be true, so we have corrected those boundary conditions by multiplying them with a correction factor45 f as per Eq. (52). This modified boundary condition takes care of the flow conditions at reactor exit in an intelligent manner. If hydrodynamic conditions at the reactor exit are such that liquid approaches the plug flow (PFR), i.e. PeL → ∞ , the correction factor f goes to 1 (Eqns. (52 and 53)), which is the same condition we would have obtained from the governing equation, by dropping the Péclet number term and interfacial mass transfer term, and then evaluating the rest of the equations at z = 1 . In this case, we simply get concentration derivatives with respect to z in terms of rate equations evaluated at the end of reactor as our boundary condition. Similarly, if 23 ACS Paragon Plus Environment


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hydrodynamic conditions correspond to a mixed flow (MFR), i.e. PeL → 0 at the end of reactor, the correction factor f goes to 0 (Eqns. (52 and 53)), and we get the conventional Dankwert’s boundary condition of zero derivative. If hydrodynamic conditions are neither corresponding to a PFR nor to a MFR, even then the boundary condition takes care of derivatives at the end of reactor accordingly, depending upon the value of Péclet number.

3.3.2

Solution algorithm for ADM The axial dispersion model and boundary conditions described so far are applicable to a

single pass co-current operation of gas and liquid only. To mimic the real experimental conditions of full liquid recycle and multiple passes through the mixer over a given batch time, an innovative solution strategy was adopted. After the first once-through solution of the model, dimensionless concentrations of all the species were noted at the end of static mixer (corresponding to one liquid residence time), and then those were taken as the inlet boundary condition for next solution cycle, owing to full recycle mode operation of liquid. The model was again solved to get species concentration corresponding to two liquid residence times, and so on. This solution process was repeated till the sum of all liquid residence times i.e.

∑τ

L

reached the

total batch time. In this manner, discrete Ce(III) conversion versus time data sets were then generated and compared (Section 4 and 5) against all the experimental conditions explored in the study. 3.4

Extension of models to other non-nuclear applications As described in Section 1, besides the nuclear decontamination, Ce(IV) has also got many

other applications for mediated oxidation of organic/inorganic pollutants in acidic media, with


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ceric regeneration by ozone. Models presented herein can be extended to those systems, which may also involve different type of gas-liquid contactors e.g. bubble columns, sieve plate columns, gas-induction reactors etc. For doing so, simply the correlations for volumetric mass transfer coefficient k L a (Eq. (15)), gas holdup ε G (Eq. (16)) and Péclet number (Eq. (41)) can be changed accordingly. In addition to that, a term for generation of Ce(III) or consumption of Ce(IV) by the target compound can be added to Eq. (28) or Eq. (44), with associated stoichiometric coefficients. In such cases, care may be taken for the coupling of mediated oxidation of substrate by Ce(IV) and the direct oxidation by ozone, if present. 4

Results and discussion All the experiments were conducted at bench scale for 2 hours (except those for 1.5 hours in

Section 4.5, due to operational constraints), with total liquid volume of 5 L. In order to keep a minimum disposable volume of used cerium solution during the experimental campaign, not all of them were repeated. To confirm the reproducibility of data, two experiment in the beginning, with ozone concentration of 29 ± 2.5 g/m3 (NTP) and 44 ± 2.5 g/m3 (NTP) in Section 4.2, were conducted twice under respective process conditions. An excellent text book match was found in both the runs, with maximum standard error of only 2.07% and 2.38% respectively, thus giving negligible error bars. Same degree of accuracy was then assumed for the further experiments. 4.1

Numerical solution of the models and simulation inputs The semi-batch mixed reactor model and the axial dispersion model were numerically

solved using explicit Runge-Kutta 4th order and Lobatto IIIa formula respectively. Effects of the various operating parameters on Ce(III) conversion with time were studied and compared with



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the experimental data. Common simulation inputs to both the models are mentioned in the captions of respective figures. Following sections describe the results so obtained. 4.2

Effect of ozone gas concentration Ozone concentration in the gas phase was varied from 10 ± 1.5 g/m3 (NTP) to 44 ± 2.5

g/m3 (NTP) by increasing the current to ozone generator, without changing the gas flow rate and other variables i.e. Ce(III) initial concentration, initial nitric acid concentration, temperature and liquid flow rate. Higher Ce(III) conversions were observed for higher gas phase concentrations. This was expected, because a higher gas phase concentration of ozone would have led to an improved solubility of gas in the solution, thereby resulting into an increased rate of reaction owing to the positive order with respect to liquid phase ozone concentration in Eq. (10). As shown in Figure 7, Ce(III) conversions of 20.0, 52.8, 57.2, and 71.6% were obtained corresponding to the ozone concentrations of 10 ± 1.5, 29 ± 2.5, 34 ± 2.0, and 44 ± 2.5 g/m3 (NTP), in 2 hours of batch time. SBM and ADM simulations matched well with the experimental data, with a mean absolute percentage deviation (MAPD) of 21.1% and 19.3%, respectively (Table 1).


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Figure 7. Effect of gas phase ozone concentration ([Ce(III)] = 0.0125 M, [HNO3] = 4 M, TC = 35.9 ± 0.5 °C, QG = 5 L/min and QL = 11 L/min).



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Table 1. Statistical comparison of experimental data with model predictions, for variation of ozone gas concentration. Parameter (G 0 )

j

(g/m3 (NTP))

10 ± 1.5

29 ± 1.5

34 ± 2.0

44 ± 2.5

Experi mental Ce(III) convers ion

Modeled Ce(III) conversion by SBM

E i (%)

M

Modeled Ce(III) conversion by ADM

MADP for SBM N

MAPDj = ∑ i=1

SBM i

(%)

M

ADM i

Ei − MiSBM  100  (%) Ei  N 

MADP for ADM N

MAPD j = ∑ i =1

E i − M iADM  100   (%) Ei  N 

(%)

5.6

3.8

4.0

12.0

8.1

8.0

16.8

12.2

11.9

20.0

16.2

15.7

17.2

12.2

13.4

31.6

25.1

25.7

42.8

36.8

36.9

52.8

47.7

47.1

21.2

13.8

15.3

35.2

28.4

29.3

48.4

41.4

41.7

57.2

53.1

52.7

28.4

18.3

20.8

49.2

37.3

38.8

60.4

53.4

54.1

71.6

67.0

66.8

27.8

27.9

18.4

16.4

19.0

16.6

19.4

16.3

21.1

19.3

N

Overall MAPD for effect of (O3 ) =

∑ MAPD

j

j =1

N


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4.3

Effect of initial concentration of Ce(III) Several experiments were conducted under different Ce(III) initial concentrations.

Concentrations of 0.0125, 0.0250, 0.05, and 0.1 M resulted in corresponding conversions of 52.8, 49.0, 37.9, and 26.2% in 2 hours. Enhanced reaction rates were observed corresponding to higher initial cerium concentrations, as indicated by relatively steeper profiles of Ce(III) concentration versus time plots in Figure 8. This can also be explained from the positive exponent of Ce(III) concentration in the rate equation (Eq. (10)). It is to be noted here that instead of plotting the conversion versus time, we have switched over to concentration versus time plots. This is because, in this set of experiments, initial concentration of Ce(III) itself was varying, which does not then suit the definition of conversion (Eq. (1)). MAPD (calculated similarly, as in Table 1) of 28.6% for SBM and of 23.7% for ADM showed a reasonable agreement.

Figure 8. Ce(III) concentration versus time plots, under various initial Ce(III) concentrations ([HNO3] = 4 M, G0 = 29 ± 1.5 g/m3 (NTP), TC = 35.7 ± 0.5 °C, QG = 5 L/min and QL = 11 L/min). 29 ACS Paragon Plus Environment


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4.4

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Effect of initial nitric acid concentration In this set of experiments, nitric acid concentration was varied. Higher nitric acid

concentrations favored the cerium oxidation, as anticipated from kinetics. Conversions of 56.8, 71.6, and 77.2% were noted against HNO3 concentrations of 2, 4, and 5.6 M in 2 hours (Figure 9). Both SBM and ADM predicted satisfactorily with respective MAPD of 19.9% and 16.7% against the experimental data.

Figure 9. Effect of initial nitric acid concentration ([Ce(III)] = 0.0125 M, G0 = 44.9 ± 1.5 g/m3 (NTP), TC = 35.3 ± 0.5 °C, QG = 5 L/min and QL = 11 L/min).

4.5

Effect of process temperature Process temperature was varied from 28.4 to 53.2 °C in these experiments, with an

accuracy of ± 0.5 °C. An increase in temperature increased the rate of reaction, thereby leading to higher Ce(III) conversions corresponding to higher temperatures, in the same batch time. As depicted in Figure 10, at the end of 1.5 hours, conversions of 48.0, 60.4, 63.6, and 71.2% were 30 ACS Paragon Plus Environment

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attained corresponding to the temperatures of 28.4, 36, 44.9, and 53.2 °C, respectively. MAPD in this case for SBM and ADM stood at 21.7% and 17.2%, respectively. It is worth mentioning here, from the standard chemistry, that an increase in temperature leads to a decrease in the solubility of a gas in liquid. This could reduce the rate of gas absorption in reacting systems. However, in our case, the temperature dependency of kinetic rate constant, i.e. 5231.4 in Eq. (11) was much stronger ( ≈ 2.15 times) than the temperature dependency of solubility of ozone, i.e. 2428 in Eq. (5). Therefore, a favorable effect of temperature rise was observed despite of the decrease in solubility with rise in temperature. It is due this reason that a high temperature of 80 °C (close to the boiling point of nitric acid) has been recommended1 for this process, which also enhances the corrosion rates of metallic components.

Figure 10. Effect of process temperature ([Ce(III)] = 0.0125 M, [HNO3] = 4 M, G0 = 44.4 ± 2.0 g/m3 (NTP), QG = 5 L/min and QL = 11 L/min).



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4.6

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Effect of liquid flow rate In order to vary the liquid recirculation rate, air flow rate to AODD pump was changed by

adjusting the pressure to regulating valve. Three experiments were then conducted at 8.2, 13.8, and 17.2 L/min of liquid flow rate through the static mixer, keeping the gas flow rate, ozone concentration, nitric acid concentration, Ce(III) initial concentration, and temperature constant. Corresponding Ce(III) conversions of 57.6, 62, and 62.4% were observed in 2 hours (Figure 11). Experimental conversion profiles tended to overlap as the liquid flow rate was increased beyond 8.2 L/min, which could be attributed to the elimination of mass transfer resistances at higher flow rates, due to increased turbulence. SBM and ADM predicted the respective conversions within 14.1% and 11.2%. These however did not respond to the changing values of QL, in the explored range of liquid flow rates. This minor deviation could be attributed to the added discrepancies, in adapting the hydrodynamic correlations (Eqns. (15, 16)) from a relatively larger (D = 6.3×10-2 m and N e = 14.3) Sulzer SMX static mixer, to a smaller equivalent one (D = 1.9×10-2 m and N e = 10.5) used in this study.


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Figure 11. Effect of liquid flow rate ([Ce(III)] = 0.0125 M, [HNO3] = 4 M, G0 = 46 ± 1.5 g/m3 (NTP), TC = 30.2 ± 0.5 °C and QG = 5 L/min).

4.7

Effect of gas flow rate In this phase, experiments were conducted at three different gas flow rates of 3, 5, and 7

L/min, for same liquid flow rate of 11 L/min. Other parameters were kept constant, as in the earlier experiments. Among those, the most important was the ozone gas concentration, which is expected to vary as oxygen flow rate through the ozone generator is varied. For this, current to the electrode assemble of ozone generator was increased at higher oxygen flow rates, to maintain the same ozone gas concentration of 46.0 ± 1.5 g/m3 (NTP) in all the three experiments. Ce(III) conversion showed a marginal increase (Figure 12) as the gas flow rate was increased. Corresponding to the said gas flow rates, respective conversions of 67.2, 68.4, and 70% were observed in 2 hours. The overlapping behavior of Ce(III) conversion profiles, as seen during the variation of liquid flow rate in earlier section, could also be explained on similar lines of 33 ACS Paragon Plus Environment


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improved hydrodynamics due to enhanced gas-liquid interfacial area, and diminishing resistances to mass transfer. SBM and ADM too predicted the similar values and behavior, with an average numerical error of 18.4% and 13.7%, respectively.

Figure 12. Effect of gas flow rate ([Ce(III)] = 0.0125 M, [HNO3] = 4 M, G0 = 46.0 ± 1.5 g/m3 (NTP), TC = 35.4 °C and QL = 11 L/min).

5

Statistical comparison of models and experimental data A parameter wise comparisons of SBM and ADM predictions with the experimental data,

i.e. on the basis of ozone concentration, Ce(III) initial concentration, nitric acid initial concentration, temperature, liquid flow rate, and gas flow have been discussed in earlier sections. In this section, we have tried to consolidate all those results for presenting an overall statistical picture of the fit.


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Parity plot in Figure 13 displays the entire data on Ce(III) conversions predicted by semibatch model, when plotted against the corresponding experimental values. The line y = x, representing the ideal 1:1 match between model and experimental data, showed a fit with R2 of 0.8084. Overall MAPD for SBM came out to be 21.1%, and the lines representing ± 25% contained most of the scatter, except a few outliers.

Figure 13. Parity plot between SBM predictions and experimental data on Ce(III) conversions, for all process parameters.

Next, the predictions from axial dispersion model were compared against the experimental data. A similar scatter in the parity plot (Figure 14), as it was for the SBM, was obtained. R2 of the ideal line y = x stood at 0.8616, and the pair of lines representing ± 25% again contained most the data points, except some outliers, also present in this case. But the



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predictions from ADM showed an improved match with experimental data, with MAPD reducing to 17.4%, as compared to 21.1% for SBM.

Figure 14. Parity plot between ADM predictions and experimental data on Ce(III) conversions, for all process parameters.

Lastly, results of both the models were plotted against each other (Figure 15), and a text book match was found, with y = x showing a R2 of 0.9926. Lines with a bound of ± 7% encapsulated the entire data, showing an excellent agreement between both the models. ADM results were found to be slightly lower, by about ≈ 3.7%, than SBM, for all the parameters.


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Figure 15. Parity plot between SBM and ADM predictions on Ce(III) conversions, for all process parameters.

6

Conclusion In nutshell, chemical oxidation of Ce(III) to Ce(IV) by ozone in nitric acid medium was

effected, using SMX type of static mixer as a gas-liquid contactor. Influence of all the process parameters was studied in detail. Higher values of kinetic parameters, i.e. temperature, initial concentration of Ce(III), nitric acid concentration, and ozone gas phase concentration were observed to be favorable for oxidation. Increase in the magnitude of hydrodynamic parameters, i.e. gas and liquid volumetric flow rates showed a corresponding increase in Ce(III) conversions up to certain optimum values, beyond which no further gain was observed owing to the elimination of mass transfer resistances.



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Data, thus generated, were compared with two different mathematical models; one considering a semi-batch mixed reactor configuration, and another by employing the concept of axial dispersion with an innovative solution algorithm to mimic the experimental loop operating in full recycle conditions. Although the empirical correlations for hydrodynamic parameters of SMX static mixer viz. k L a , a , ε G , Pe L and the equations for ozone solubility in this study were adapted from literature for similar systems, a reasonable match between the experimental data and the predictions from both the models was obtained, with an overall average deviation of 21.1% and 17.4%, respectively. However, there exists a scope for further improvement in the predictive capabilities of models by using the equations experimentally determined for actual system at hand. These models for ozone based oxidation of Ce(III) to Ce(IV) could be useful in the design and scale up of nuclear decontamination process, after incorporating the Ce(IV) consumption by contaminated metallic components of given surface area, which will be the subject of our future publication. Same models could also be extended to other non-nuclear environmental applications, by adding a term for the rate of consumption of Ce(IV) or generation of Ce(III) ions by organic/inorganic pollutants, in the mass balance for cerium.

Acknowledgements Authors are thankful to the reviewers for their constructive comments, which have led to significant improvements in this manuscript. One of the authors (S.S.) is grateful to Shri B.V. Shah, former Head-EDS of PSDD, for his valuable technical discussions and advice. Motivation and support received from seniors at BARC, namely Smt. Smitha Manohar, Head-PEDS, Dr. J. G. Shah, Head-PSDD and Shri Kailash Agarwal, Associate Director-NRG is also acknowledged.


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Thanks are also due to other departmental colleagues including Shri A. Sahu, Shri V. Sinha, Shri R.D. Silva, Shri R.G. Gaygwal, Shri R.N. Rath and Shri K.K. Suman for their kind help in carrying out this work.

Supporting Information Pre-calculated value of mass transfer enhancement factor E for assumption (d); Calculation of ozone partial pressure from analyzer reading at NTP.

Author Information *Corresponding author: Sukhdeep Singh, Email: [email protected], Phone: +91-22-25591431, Fax: +91-22-25505185, ORCID: 0000-0002-4046-5879. Notes: Authors declare no competing financial interest.

Nomenclature A

total gas-liquid interfacial area (m2)

a

interfacial area per unit volume of contactor (m2/m3)

CA

concentration of A in liquid phase (kmole/m3)

CB

concentration of B in liquid phase (kmole/m3)

CC

concentration of C in liquid phase (kmole/m3)

cA

dimensionless concentration of A in liquid phase

cB

dimensionless concentration of B in liquid phase 39 ACS Paragon Plus Environment


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cC

dimensionless concentration of C in liquid phase

C AS

solubility of A in water (kmole/m3)

C B0

initial concentration of B in liquid phase (kmole/m3)

CC0

initial concentration of C in liquid phase (kmole/m3)

D

inner diameter of static mixer (m)

DA

diffusivity of A in water (m2/s)

D Air

diffusivity of air in water (m2/s)

DB

diffusivity of B in water (m2/s)

DL

liquid phase dispersion coefficient (m2/s)

D O2

diffusivity of oxygen in water (m2/s)

D aA

Damköhler number for oxidation reaction of A

D a Ad

Damköhler number for self-decomposition of A

D aB

Damköhler number for B

DaC

Damköhler number for C

E

mass transfer enhancement factor

E∞

mass transfer enhancement factor for an infinitely fast reaction

Ei

defined in Table 1 (%)

f

boundary condition correction factor


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G0

ozone concentration at NTP, as given by analyzer (g/m3)

GA

ozone concentration at experimental conditions (g/m3)

H

Henry’s constant for ozone (atm.mole/mole)

k

rate constant for Ce(III) oxidation reaction (m4.29/kmole1.43.s)

km,n

rate constant of an (m, n)th order reaction (m3(-1+m+n)/kmole (m+n-1).s)

kL

liquid film mass transfer coefficient (m/s)

kLa

volumetric mass transfer coefficient (1/s)

Kw

ion product constant for water (kmole2/m6)

l

length coordinate of static mixer (m)

L

length of static mixer (m)

m

order of Ce(III) oxidation reaction with respect to B

mA

mass of ozone (g)

MA

molecular weight of ozone (48 g/mol)

MC

molecular weight of nitric acid (= 63.01 g/mol)

M

molecular weight of oxygen (= 32 g/mol)

O2

M Air

molecular weight of air (= 29 g/mol)

Mi

defined in Table 1 (%)

n

order of Ce(III) oxidation reaction with respect to A

N

number of experimental data points (= 4)

nA

moles of A in liquid (kmole) 41 ACS Paragon Plus Environment


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NA

z =0

flux of A at gas-liquid interface (kmole/m2.s)

Ne

length to diameter ratio of static mixer

p

order of Ce(III) oxidation reaction with respect to C

pA

partial pressure of A in the gas phase (atm)

Pe L

Péclet number for liquid phase

pH

potential of hydrogen of the solution

QG

gas volumetric flow rate (L/min)

QL

liquid volumetric flow rate (L/min)

− rA

rate of consumption of dissolved ozone by Ce(III) (kmole/m3.s)

− rAd

rate of ozone self-decomposition reaction (kmole/m3.s)

S

cross section area of static mixer (m2)

StL

Stanton number for liquid phase

t

time in seconds (s)

uG

superficial gas velocity (m/s)

uL

superficial liquid velocity (m/s)

V

volume of gas in ideal gas equation (m3)

VAir

molal volume of air at normal boiling point (=29.9 cm3/mol)

VC

volume of contactor (m3)

VG

volume of gas (m3) 42 ACS Paragon Plus Environment

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VL

volume of liquid (m3)

T

temperature in Kelvin (K)

TC

temperature in degree Celsius (°C)

x

defined in text

z

dimensionless static mixer length coordinate

Greek letters

φ

Hatta number

νB

stoichiometric coefficient of B in oxidation reaction (= 2)

εG

gas holdup

εL

liquid holdup

ρ

defined in text

µ

defined in text

ρL

liquid density (kg/m3)

ρG

gas density (kg/m3)

µL

liquid viscosity (kg/m.s)

µG

gas viscosity (kg/m.s)

σ

association parameter for water (= 2.6)

λ

defined in text

C

defined in text



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Special symbols [ ]

species concentration in liquid (kmole/m3)

Subscripts i

ith data point in Table 1 or ith species

j

jth parameter value Table 1

0

initial value of concentration

G

gas

L

liquid

A

ozone

B

Ce(III)

C

nitric acid

Air-Water

air and water system

Abbreviations NHE

normal hydrogen electrode

TFT

thin film transistor

LCD

liquid crystal display

PHWR

pressurized heavy water reactor

AISI

American iron and steel institute

PUREX

plutonium uranium redox extraction


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TBP

tributyl phosphate

TEHDGA

tetra (2-ethylhexyl diglycolamide)

PVDF

polyvinylidene fluoride

NTP

normal temperature pressure

PPM

parts per million

RTD

resistance temperature detector

NB

nominal bore

PVC

polyvinyl chloride

SBM

semi-batch model

ADM

axial dispersion model

MAPD

mean average percentage deviation

EXP

experimental

AODD

air operated double diaphragm

References (1) Ponnet, M.; Klein, M.; Rahier, A. Chemical decontamination MEDOC® using cerium(IV) and ozone. WM’00 Conference, Tucson, AZ. 2000. (2) Shah, J. G.; Dhami, P. S.; Gandhi, P. M.; Wattal, P. K. Decontamination of alpha contaminated metallic waste by cerium(IV) redox process. Bhabha Atomic Research Center (BARC)-News Letter. 2012, 328, 56-61. (3) Caire, J. P; Cullie, S.; Dalard, F.; Fulconis, J. M.; Delagrange, H. AISI 304 L stainless steel decontamination by a corrosion process using cerium(IV) regenerated by ozone Part I: Study of the accelerated corrosion process. J. Applied Electro. 2003. 33, 703-708.



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(4) Singh, S.; Bhaskarwar, A. N.; Vincent, T.; Wattal, P. K. Kinetics of oxidation of Ce(III) to Ce(IV) by ozone in nitric acid medium. J. of Radioanal. and Nucl. Chem. 2016, 309, 691-700. (5) Mizuno, T.; Tsuno, H.; Yamada, H. Development of ozone self-decomposition model for engineering design. Ozone: Sci. & Engg. 2007, 29, 55-63. (6) Pohjanne, P.; Sitala, M. In-Plant Corrosion Testing in Ozone Bleaching Environments. http://www.outokumpu.com/SiteCollectionDocuments/Outokumpu-Acom-In-plantcorrosion-testing-in-ozone.pdf (accessed June 08, 2017). (7) Arenas, L. F.; Leon, P. C.; Walsh, F. C. Electrochemical redox processes involving soluble cerium species. Electrochimica Acta. 2016, 205, 226-247. (8) Binnemans, K. Applications of tetravalent cerium compounds. In Handbook on the Physics and Chemistry of Rare Earths; Gschneidner, K. A., Bunzli, J. C., Pecharsky, V. K., Eds.; Elsevier: North Holland, 2006; Vol. 36, pp 281-392. (9) Kramer, K.; Robertson, P. M.; Ibl, N. Indirect electrolytic oxidation of some aromatic derivatives. J. Appl. Electrochem. 1980, 10, 29-36. (10) Singh, M. P.; Singh, H. S.; Verma, M. K. Kinetics and mechanism of the ruthenium(1 1 1) chloride catalyzed oxidation of butanone-2 and pentanone-3 by cerium(IV) sulfate in aqueous sulfuric acid medium. J. Phys. Chem. 1980, 84, 256-259. (11) Ruoff, P.; Nevdal, G. Kinetics and mechanism of the Ce(IV) oxidation of methylmalonic acid in H2SO4. J. Phys. Chem. 1989, 93, 7802-7806. (12) Cho, L. Y.; Madurro, J. M.; Romero, J. R. Electrooxidation of β -dicarbonyl compounds using ceric methanesulfonate as mediator: some kinetics and spectroscopic studies. J. Catal. 1999, 186, 31-35.


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(21) Matheswaran, M.; Balaji, S.; Chung, S. J.; Moon, I. S. Mediated electrochemical oxidation of phenol in continuous feeding mode using Ag (II) and Ce (IV) mediator ions in nitric acid: a comparative study. Chem. Engg. J. 2008, 144, 28-34. (22) Balaji, S.; Chung, S. J.; Matheswaran, M. Destruction of organic pollutants by cerium(IV) MEO process: a study on the influence of process conditions for EDTA mineralization. J. Hazard. Mater. 2008, 150, 596-603. (23) Balaji, S.; Chung, S. J.; Ryu, J. Y.; Moon, I. S. Destruction of commercial pesticides by cerium redox couple mediated electrochemical oxidation process in continuous feed mode. J. Hazard. Mater. 2009, 172, 1470-1475. (24) Huang, K. L.; Chen, T. S.; Yeh, K. J. C. Regeneration of Ce(IV) in simulated spent Cretching solutions using an undivided cell. J. Hazard. Mater. 2009, 171, 755-760. (25) Chen, T. S.; Yeh, K. J. C.; Huang, K. L. Anion effects on the electrochemical regeneration of Ce(IV) in nitric acid used for etching chromium. J. Hazard. Mater. 2008, 152, 922-928. (26) Nzikou, M.; Aurousseau, M.; Apicque, F. Electrochemical investigations of the Ce(III)/CeIV) couple related to a Ce(IV)-assisted process for SO2/NOx abatement. J. Appl. Electrochem. 1995, 25, 967-972. (27) Pillai, C. K.; Raju, T.; Chung, S. J.; Moon, I. S.; Removal of H2S using a new Ce(IV) redox mediator by a mediated electrochemical oxidation process. J. Chem. Technol. Biotechnol. 2009, 84, 447-453. (28) Fujita, R.; Enda, M.; Morisue, T. Redox decontamination technique development, (I). J. of Nuclear Sci. and Tech. 1989, 26, 339-349.


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Oxidation of Ce(III) to Ce(IV) by Ozone in Nitric Acid Medium Using a

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