Thermally Safe Operation of Nitric Acid Oxidation

Runaway and thermally safe operation
of a nitric acid oxidation in a
semi-batch reactor
B.A.A. van Woezik

RUNAWAY AND THERMALLY SAFE OPERATION
OF A NITRIC ACID OXIDATION IN A
SEMI-BATCH REACTOR
PROEFSCHRIFT
ter verkrijging van
de graad van doctor aan de Universiteit Twente,
op gezag van de rector magnificus,
prof. dr. F.A. van Vught,
volgens besluit van het College voor Promoties
in het openbaar te verdedigen
op vrijdag 22 september 2000 te 13.15 uur.
door
Bob Arnold August van Woezik
geboren op 6 januari 1969
te Nijmegen

Dit proefschrift is goedgekeurd door de promotor
Prof.dr.ir. K.R. Westerterp

This research was supported by the Technology Foundation STW, applied
science division of NWO and the technology program of the Ministry of
Economic Affairs.
Copyright © 2000 B.A.A. van Woezik, Eindhoven, The Netherlands
No part of this book may be reproduced in any form by any means, nor
transmitted, nor translated into a machine language without written permission
from the author.
CIP-GEGEVENS KONINKLIJKE BIBLIOTHEEK, DEN HAAG
Woezik, Bob Arnold August van
Runaway and thermally safe operation of a nitric acid oxidation in a semi-batch
reactor / Bob Arnold August van Woezik.
Thesis University of Twente, Enschede. – With ref. – With summary in Dutch.
ISBN 90 - 365 14878
Subject headings: runaway, liquid-liquid reactions, nitric acid oxidation.

1
Summary and Conclusions
A number of serious accidents has occurred due to a runaway reaction of a
heterogeneous liquid-liquid reaction whereby a secondary side reaction was
triggered. A basic lack of proper knowledge of all the phenomena, occurring in
such a system, is one of the prime causes that may lead to overheating and
eventually a thermal runaway. Therefore, a better understanding of these kinds
of processes is of great importance for the safe and economic design as well as
safe operation of those reactions. This thesis deals with the safe operation of a
multiple liquid-liquid reaction in a semi-batch reactor in the example of the
nitric acid oxidation of 2-octanol. A general introduction about runaways in
(semi) batch reactors is given in Chapter 1.
In Chapter 2 the oxidation of 2-octanol with nitric acid is studied. The oxidation
of 2-octanol with nitric acid has been selected as a model reaction for a
heterogeneous liquid-liquid reaction with an undesired side reaction. 2-Octanol
is first oxidized to 2-octanone, which can be further oxidized to carboxylic
acids. The oxidation of 2-octanol and 2-octanone with nitric acid exhibits the
typical features of nitric acid oxidations, like a long induction time without
initiator; autocatalytic reaction; strong dependence of mineral acid concentration
and high energy of activation. However, there is a limited knowledge of the
exact chemical structure of the compounds in the aqueous reaction phase and of
a number of unknown, unstable compounds in the organic phase. Next to this
the exact mechanism is still not elucidated. As a consequence of this, a
considerable model reduction was necessary to describe the overall reaction
rates.
An extensive experimental program has been followed using heat flow
calorimetry supported by chemical analysis. The oxidation reactions have been
carried out in a reaction calorimeter RC1 of Mettler Toledo, which contains a
jacketed 1-liter glass vessel. The reactions have been studied in the range 0 to 40
ºC, with initial nitric acid concentrations of 50 to 65 wt% and a stirring rate of
700 rpm. The kinetic constants have been determined for both reactions. The
observed conversion rates of the complex reactions of 2-octanol and 2-octanone
with nitric acids can be correlated using only two kinetic equations, in which the
effect on temperature is described through the Arrhenius equation and the effect
on acid strength through Hammett’s acidity function.

2
The nitric acid and the organic solution are immiscible, so chemical reaction and
mass transfer phenomena occur simultaneously. The results indicate the
oxidation of 2-octanol is operated in the non-enhanced regime when nitric acid
is below 60 wt% or when the temperature is below 25 ºC at 60 wt% HNO3,
while the oxidation of 2-octanone is operated in the non-enhanced regime for the
whole range of experimental conditions considered. Under these conditions the
mass transfer resistance does not influence the overall conversion rate, so the
governing parameters are the reaction rate constant and the solubility of the
organic compounds in the nitric acid solution. This has also been experimentally
confirmed by determining the influence on stirring rate.
In parallel a model has been developed to describe the conversion rates, that
successfully can predict the behavior of the semi-batch reactor, i.e. concentration
and temperature time profiles. The experimental results and simulations are in
good agreement and it has been found possible to describe the thermal behavior
of the semi-batch reactor for the nitric acids oxidation reactions with the film
model for slow liquid-liquid reactions and a simplified reaction scheme.
In Chapter 3 the thermal behavior of this consecutive heterogeneous liquid-
liquid reaction system is studied in more detail by experiments and model
calculations. An experimental installation has been built, containing a 1-liter
glass reactor, followed by a thermal characterization of the equipment. Two
separate cooling circuits have been installed to study different cooling
capacities: a cooling jacket and a cooling coil. The reactor has been operated in
the semi-batch mode under isoperibolic conditions, i.e. with a constant cooling
temperature. A series of oxidation experiments has been carried out to study the
influence of different initial and operating conditions. The thermal behavior has
been studied with a coolant temperature of -5 to 60 ºC, a dosing time of 0.5 to 4
hours, an initial nitric acid concentration of 60 wt% and a stirring rate of 1000
rpm.
The reaction is executed in a cooled SBR in which the aqueous nitric acid is
present right from the start and the organic component 2-octanol is added at a
constant feed rate. The 2-octanol reacts to 2-octanone, which can be further
oxidized to unwanted carboxylic acids. A dangerous situation may arise when
the transition of the reaction towards acids takes place in such a fast way that the
reaction heat is liberated in a very short time and it results in a temperature
runaway. The use of a longer dosing time or a larger cooling capacity effectively
moderates the temperature effects and it will eventually even avoid such an
undesired temperature overshoot. In the later, the process is regarded as
invariably safe and no runaway will take place for any coolant temperature and

3
the reactor temperature will always be maintained between well-known limits.
The conditions leading to an invariably safe process are determined
experimentally and by model calculations.
Because of the plant economics one must achieve a high yield in a short time
and under safe conditions. The reaction conditions should rapidly lead to the
maximum yield of intermediate product 2-octanone and after that the reaction
should be stopped at the optimum reaction time. The appropriate moment in
time to stop the reaction can be determined by model calculations. The influence
of operation conditions, e.g. dosing time and coolant temperature, on the
maximum yield are studied and will be discussed.
In the oxidation of 2-octanol one focuses on the first reaction because high
yields of ketone are required, while the danger of a runaway reaction must be
attributed to the ignition of the secondary reaction. The reaction system can be
considered as two single reactions and, therefore, also the boundary diagram
− developed by Steensma and Westerterp [1990] − for single reactions has been
used to estimate critical conditions for the multiple reaction system. The
boundary diagram can be used to determine the dosing time and coolant
temperature required for safe execution of the desired reaction. However, for
suppression of the undesired reaction it leads to too optimistic coolant
temperatures.
Studying the dynamic behavior of heterogeneous liquid-liquid reactions involves
a number of difficulties, because chemical reaction and mass transfer
phenomena occur simultaneously. The interfacial area is essential for an
accurate prediction of the mass transfer and chemical reaction rates in liquid-
liquid reactions. The interfacial area for a liquid-liquid system in a mechanically
agitated reactor is determined in Chapter 4. This has been done by means of the
chemical reaction method. This method deals with absorption accompanied by a
fast pseudo-first order reaction. The saponification of butyl formate ester with 8
M sodium hydroxide solution has been used. The extraction rate is determined
in a stirred cell with a well-defined interfacial area equal to 33.4 cm2
and a
correlation has been derived to describe the mole flux of ester through the
interface. The kinetic rate constants have been calculated and are compared to
data from literature. The reaction is affected by the amount of ions in the
solution. The reaction rate constant is described by an extra term in the usual
Arrhenius equation to account for this effect of the ionic strength.
The reactor, with a total volume of 0.5 liter, has been operated continuously to
study the interfacial area in a turbulently mixed dispersion. A correlation has

4
been derived for the Sauter mean diameter for both, reaction in the dispersed
phase as well as reaction in the continuous phase. A viscosity factor had to be
incorporated to obtain one single correlation. The Sauter mean diameter can be
described by correlations similar to those in literature, only the constants
deviate, because the specific properties of the system investigated and the
reactor configuration are different. These constants were found to depend also
on the phase that is dispersed. With the organic ester phase dispersed, droplet
diameters were found between 35 and 75 µm and between 65 and 135 µm in
case the aqueous phase is dispersed. The drop size seems to be influenced by the
density of the continuous phase as well as the ratio of the viscosities of the two
phases. It is not unambiguous which phase dispersed will give the smallest drop
size and, hence, the largest interfacial area. It is, therefore, recommended to
determine the drop size for both liquids as the dispersed phase.
The mass transfer with reaction is described using the film theory. This model
can usually be applied within the uncertainties of the estimated physico-
chemical parameters, even though it is the simplest approach. The validation for
the chemically enhanced reaction regime is presented. The necessary conditions
are all full-filled in all experiments except that of a large Hinterland ratio.
Therefore, the reaction between ester and sodium hydroxide in a single drop has
been described numerically. The effect of a small Hinterland ratio shows itself
by the inability of either the film theory or penetration theory to allow for
eventual depletion of the reactant within the droplet. For the used experimental
set-up and experimental conditions, the contact time is relatively short and
deviations due to depletion of NaOH in the droplet are not to be expected. For
the smallest experimentally determined droplet diameters, the assumption of a
flat interface is no longer valid and the influence of the curvature of the interface
has to be taken into account, otherwise the film theory can be used with
confidence.
References
Steensma, M. and Westerterp, K.R., Thermally safe operation of a semi-batch
reactor for liquid-liquid reactions. Slow reactions, Ind. Eng. Chem. Res. 29
(1990) 1259-1270.

5
Contents
Summary and Conclusions 1
Chapter 1: General Introduction 9
1.1 General 11
1.2 Present work 13
References 14
Chapter 2: The nitric acid oxidation of 2-octanol and 2-octanone 17
Abstract 18
2.1 Introduction 19
2.2 Oxidation reactions with nitric acid 19
Oxidation of 2-octanol
Oxidation of 2-octanone
2.3 Derivation of overall conversion rates 22
Kinetic expressions
Conversion rates in a semi-batch reactor
2.4 Experimental set-up and principle of measurements 27
Reaction calorimeter
Experimental set-up and experimental procedure
Chemical treatment and chemical analysis
2.5 Experimental results 34
Identification of reaction regime
Determination of kinetic parameters
2.6 Simulation of isothermal runs 45
2.7 Model validation and limitations 49
Model verification with isoperibolic experiments
2.8 Discussion and conclusions 55
Notation 56
References 59

Contents
6
Chapter 3: Runaway behavior and thermally safe operation of multiple
liquid-liquid reactions in the semi-batch reactor 63
Abstract 64
3.1 Introduction 65
3.2 Nitric acid oxidation in a semi-batch reactor 66
Reaction system
Mathematical model
3.3 Thermal behavior of the nitric acid oxidation of 2-octanol 75
Sudden reaction transition
Gradual reaction transition
3.4 Recognition of a dangerous state 86
3.5 Experimental set-up and procedure 88
Thermal characterization of equipment
Check on the validity of the model for slow reactions
3.6 Experimental results 95
Temperature profiles
Thermally safe operation of the nitric acid oxidation
Influence of dosing time
Influence of cooling capacity
Invariably safe operation
3.7 Prediction of safe operation based on the individual reactions 105
Notation 109
References 112
Chapter 4: Determination of interfacial areas with the chemical
method for a system with alternating dispersed phases 113
Abstract 114
4.1 Introduction 115
4.2 Measurement of interfacial area, the theory 116
Determination by the chemical method
4.3 Experimental set-up 120
4.4 Measurements in the stirred cell 123
Experimental procedure
Determination of flux equation
Calculation of kinetics

Contents
7
4.5 Determination of interfacial area 130
Experimental procedure
Determination of drop size correlation
4.6 Validity of the assumed conditions 137
The effect of small Hinterland ratio
Notation 146
References 148
Appendix 4.A: Physico-chemical parameters 151
Appendix 4.B: Numerical model 154
Samenvatting en conclusies 155
Dankwoord 159
List of publications 162
Levensloop 163

General Introduction
11
Temperature
Heatrates
Heat production rate
Heat removal rate 2
1
1.1 General
At Seveso on July 10th
1976 a runaway reaction took place that led to a
discharge of highly toxic dioxin contaminating the neighboring village. The
runaway reaction in the unstirred mixture took place seven hours after stirring
had been stopped and was triggered by a small heat input from the hot wall, see
Kletz [1988]. It turned out to be one of the best-known chemical plant accidents
and it became clear that the safety margins had not been recognized. The
accident induced the fine chemicals industry to review their safety systems and
to develop more refined methods for safeguarding their reactors.
A considerable number of accidents has occurred, that can be attributed to this
so-called runaway reaction. The basic understanding of a runaway reaction
arises from the thermal explosion theory according to Semenov. This theory
deals with the competition between heat generation by an exothermic reaction
and heat removal from the reaction mass to, for instance, the cooling jacket. The
heat generation depends, according to Arrhenius, exponentially on temperature,
while the heat removal depends linearly on temperature, see Figure 1.
Figure 1: Heat flow diagram. Heat production rate by chemical reaction and
heat removal rate by cooling.

Chapter 1
12
A steady state will be reached as soon as the heat production rate is equal to the
heat removal rate. This will be the case for both the temperatures of the
intersections in Figure 1. The degree of control of the heat production rate
directly follows from this plot. At intersection (1) the slope of the heat removal
line is greater than that of the heat production curve and consequently a small
deviation from this steady state automatically results in a return to its origin.
Therefore, intersection (1) represents a stable operation point and the exothermic
reaction is under control. On the other hand, intersection (2) represents an
unstable operation point. If, for some reason, a temperature deviation occurs, the
original operating conditions will never be reached again. In case of a
temperature decrease the steady state of intersection (1) will be attained. In case
of an increase, the rate of heat generation will always exceed that of the heat
removal. This will lead to an unhindered self-acceleration of the reaction rate
and thereby of the heat production rate, which is known as a runaway reaction.
When the reaction is carried out in the batch reactor the process will not reach a
steady state. The batch reactor has great flexibility and is therefore extensively
used in the production of fine and specialty chemicals and accordingly
contributes to a significant part of the world’s chemical production in number
and value. However, batch processes are usually very complex with strong non-
linear dynamics and time-varying parameters. The process requires a continuous
safeguarding and correction by the operator. Furthermore, due to the small
amounts produced and variety of processes, obtaining complete understanding
of the reactor dynamics is usually not economically feasible. This lack of
knowledge gave rise to a number of accidents. Barton and Nolan [1991] have
reported the prime causes of industrial incidents, which were mainly related to
the lack of knowledge of the process chemistry, to inadequate design and to
deviation from normal operating procedures. The study of accidents also shows
that batch units are usually more frequently involved in accidents than
continuous process plants.
An attractive way to reduce the potential hazard is to avoid the use of truly batch
reactions and instead switch to semi-batch. With this type of operation the
reactor is initially charged with one of the reactants and the other reactants are
added continuously to the vessel. This makes it possible to control the reaction
rate and hence the generation of heat. Therefore, semi-batch reactors are often
used for highly exothermic reactions.
For semi-batch reactors with homogeneous reaction systems Steinbach [1985]
and Hugo and Steinbach [1985] demonstrated that too low reaction temperatures
could cause runaways. If the initial temperature is too low, the added reactants

13
will not react immediately and will start to accumulate. Under certain
circumstances the combination of increasing concentration and a gradual
temperature rise may lead to a runaway. Criteria for safe operation of a semi-
batch reactor are based on the prevention of accumulation of unreacted
reactants. The semi-batch reactor should therefore be operated with a
temperature high enough to maintain the reaction rate approximately equal to
the feed rate.
A great number of industrial processes in semi-batch reactors involve systems in
which two immiscible phases coexist, generally an organic and an aqueous one.
Like in the manufacturing of organic peroxides, sulphonates, nitrate esters and
other nitrocompounds. Steensma and Westerterp [1990, 1991] developed models
for liquid-liquid reactions to study thermal runaways taking place in such
heterogeneous systems. In case the reaction takes place in the dispersed phase,
the system was found to be more prone to accumulation than when the reaction
takes place in the continuous phase. In the latter case, the system exhibits a
better conversion rate at the start, which reduces the danger of runaway
reactions. Also a distinction could be made between slow reactions, where the
reaction takes place in the bulk of one of the liquid phases, and fast reactions -
i.e. chemical enhanced - with reaction in the boundary layer of one of the
phases. A runaway can occur in liquid-liquid reaction systems due to
accumulation of the added reactants in the reacting phase for slow reactions, and
in the non-reacting phase for fast reactions.
Although the contents of a reactor vessel may normally yield the desired
reaction products, deviations from normal operating conditions or upset
conditions such as loss of jacket cooling can lead to increased temperatures.
This may initialize unwanted decomposition reactions, elevate the system
pressure and lead to an emission as in the case of Seveso. The general approach
in preventing a runaway reaction is to avoid triggering off side and chain
reactions. It is a rather conservative approach, while in some cases it is
inevitable to allow an unwanted reaction partially to take place.
1.2 Present work
The thermal behavior is studied of a multiple liquid-liquid reaction in a semi-
batch reactor. The main goal is to understand and to ensure safe operation of this
kind of system by means of experiments and model calculations.

Chapter 1
14
Experimental studies of the thermal behavior of runaway reactions in a (semi)
batch reactor are scarce and no experimental systems have been described in
detail in which strongly exothermic side reactions can be triggered. The
oxidation reaction of 2-octanol has been chosen as a model reaction. Chapter 2
deals with the kinetic study of the nitric acid oxidation of 2-octanol to 2-
octanone and to the further oxidation products. The reactions have been studied
in a reaction calorimeter and a model, based on the film theory, has been
developed to describe the conversion rates.
In chapter 3 the nitric acid oxidation of 2-octanol is used to study experimentally
the thermal runaway behavior of an exothermic heterogeneous multiple reaction
system in a 1-liter glass reactor. The reactor is operated in a semi-batch manner
with a constant cooling temperature. Typical reaction regions can be
distinguished with increasing operation temperatures, which will be
demonstrated and explained. Parameters are studied to produce the required
intermediate product, 2-octanone, with a high yield and in a safe manner. The
results of the simulations are compared to the experimental observations.
One of the causes of accidents, see Barton [1991], is that the phenomena in, for
instance, liquid-liquid reactions are not understood. Essential for an accurate
prediction of the mass transfer and chemical reaction rates in liquid-liquid
reactions is the interfacial area. Chapter 4 deals with the interfacial area in a
mechanically agitated reactor. The interfacial area of a liquid-liquid system has
been determined by the chemical reaction method using the saponification of
butyl formate ester. Although drop sizes in dispersions have been studied
extensively, experimental data for the same system and alternating phases
dispersed are scarce. In this chapter the results are given for the two types of
dispersion. The mass transfer with reaction is described using the film theory
and the necessary conditions are verified. For the smallest droplets with hardly
any bulk, the film model is not realistic anymore. Induced deviations are studied
and discussed.
References
Barton, J.A. and Nolan, P.F., Incidents in the chemical industry due to thermal-
runaway chemical reactions. In: Euro courses, Reliability and risk analysis,
Vol.1: Safety of Chemical Batch Reactors and Storage Tanks, A. Benuzzi and
J.M. Zaldivar (eds.), Kluwer Academic, Dordrecht 1991, pp. 1-17.

15
Hugo, P. and Steinbach J., Praxisorientierte Darstellung der thermischen
Sicherheitsgrenzen für den indirekt gekühlten Semibatch-Reaktor. Chem. Ing.
Tech. 57 (1985) 780-782.
Kletz, T., Learning from accidents in industry, Butterworths, London 1988, pp.
79-83.
Steensma, M. and Westerterp, K.R., Thermally safe operation of a semibatch
reactor for liquid-liquid reactions. Slow reactions, Ind. Eng. Chem. Res. 29
(1990) 1259-1270.
Steensma, M. and Westerterp, K.R., Thermally safe operation of a semibatch
reactor for liquid-liquid reactions - Fast reactions, Chem. Eng. Technol. 14
(1991) 367-375.
Steinbach, J., Untersuchung zur thermischen Sicherheit des indirekt gekühlten
Semibatch-Reaktors, PhD-thesis, Technical University of Berlin, Berlin, 1985.

2
The Nitric Acid Oxidation of
2-Octanol and 2-Octanone

Chapter 2
18
Abstract
The oxidation of 2-octanol with nitric acid has been selected as a model reaction
for a heterogeneous liquid-liquid reaction with an undesired side reaction. 2-
Octanol is first oxidized to 2-octanone, which can be further oxidized to
carboxylic acids. An extensive experimental program has been followed using
heat flow calorimetry supported by chemical analysis. A series of oxidation
experiments has been carried out to study the influence of different initial and
operating conditions such as temperature, stirring speed and feed rate. In parallel
a semi-empirical model has been developed to describe the conversion rates.

The Nitric Acid Oxidation of 2-Octanol and 2-Octanone
19
2.1 Introduction
A number of incidents concerning runaway reactions involve systems in which
two immiscible phases coexist, generally an organic and an aqueous one.
Examples of such systems, in which simultaneously mass transfer and chemical
reaction are important, are nitrations, sulphonations, hydrolyses, esterifications
and oxidations. Experimental studies of the thermal behavior of runaway
reactions in a (semi) batch reactor are scarce. Only homogeneous reaction
systems are described in literature: the homogeneous, sulfuric acid catalyzed
hydrolysis of acetic anhydride, see e.g. Haldar and Rao [1992a,b] and the
homogeneous, acid catalyzed esterification of 2-butanol and propionic
anhydride, see Snee and Hare [1992]. No experimental systems have been
described in detail for a heterogeneous liquid-liquid reaction, in which strongly
exothermic side reactions can be triggered. However, in many nitrations it is
known that dangerous side reactions can play a role like undesired oxidation
reactions, see Camera et al. [1983]. They studied the oxidation of ethanol with
nitric acid, where decomposition reactions can give rise to explosions.
To study the thermal behavior of a liquid-liquid reaction the oxidation of a long
chain alcohol with nitric acid has been chosen. The ketones formed in the
oxidation of secondary alcohols are more stable than aldehydes, so the oxidation
of 2-octanol with nitric acid has been chosen as a model reaction. Secondary
alcohols are also oxidized in the commercial production of adipic acid, in which
cyclohexanol is oxidized. This reaction has been studied by van Asselt and van
Krevelen [1963a,b,c,d] and has been reviewed by Castellan et al. [1991].
This work presents experimental data for the oxidation of 2-octanol to 2-
octanone and further oxidation products. The main objective is to develop a
model to describe the conversion rates of 2-octanol and 2-octanone.
2.2 Oxidation reactions with nitric acid
Nitric acid is a commonly used oxidizer. Especially alcohols, ketones, and
aldehydes are oxidized to produce the corresponding carboxylic acids, for
instance adipic acid, see Davis [1985]. The oxidation of cyclohexanol with nitric
acid is very similar to the oxidation of 2-octanol, see Castellan et al. [1991]. The
mechanism of these nitric acid oxidations is still not elucidated. Oxidations with
nitric acid are in general very complex and usually several intermediates are
formed, see e.g. Ogata [1978]. The elucidation of the real pathways was beyond

Chapter 2
20
the scope of the project: therefore, it has been chosen to simplify the description
of the conversion rates of 2-octanol and 2-octanone.
The oxidation of 2-octanol occurs in a two-phase reaction system in which a
liquid organic phase, containing 2-octanol, is contacted with an aqueous, nitric
acid phase. The main organic components during the reactions can be
represented as follows:
These reactions are further described in more detail in the following paragraphs.
Experimental results of nitric acid oxidations from literature will also be used.
Oxidation of 2-octanol
Different reacting species have been proposed like N2O4 by Horvath et al.
[1988], NO+
by Strojny et al. [1971] and NO2 by Camera et al. [1983]. Castellan
et al. [1991] concluded that at ambient temperatures the oxidation proceeds
mainly via an ionic-molecular mechanism. This indicates that the (NO+
)
nitrosonium ion mechanism is applicable for the conditions used in this work.
This ion can be formed from nitrous acid and nitric acid through reaction (1):
HNO HNO NO NO H O2 3 3 2+ ↔ + ++ −
(1)
The oxidations with pure nitric acid exhibit in general a long induction period,
see e.g. van Asselt and van Krevelen [1963a] and Ogata et al [1966]. This
induction time can be shortened or even eliminated by adding an initiator like
NaNO2, which forms nitrous acid:
NaNO H O HNO Na H O2 3 2 2+ → + ++ +
(2)
The reaction is completely suppressed by addition of urea, which reacts with
nitrous acid, see e.g. Camera et al. [1979], according to:
2 2 32 2 2 2 2 2HNO CO NH N CO H O+ → + +( ) (3)
This is in agreement with the above-mentioned formation of a nitrosonium ion
or its equivalent.
2-octanone2-octanol carboxylic acids

21
Figure 1: Reaction pathway for the oxidation of 2-octanol with nitric acid.
R = CH3(CH2)4—
The oxidation of 2-octanol to 2-octanone proceeds via the formation of an
intermediate, which has been identified, as 2-octyl nitrite, using GC-MS. The
reaction pathway of the first steps of the oxidation of 2-octanol can be
schematically represented as in Figure 1. After addition of the initiator, HNO2 is
formed, the oxidation starts and proceeds autocatalytically. One molecule of
HNO2 - or NO+
according to Equation (1) - is consumed in the first step, while
two are formed in the second step. This net formation of an equimolar amount of
HNO2 also has been found for the oxidation of cyclohexanol to cyclohexanone,
see van Asselt and van Krevelen [1963a, d].
Oxidation of 2-octanone
2-Octanone can be further oxidized to carboxylic acids. During this reaction an
equimolar amount of nitrous acid is consumed, the same as in the oxidation of
cyclohexanone, see van Asselt and van Krevelen [1963a].
Figure 2: Reaction pathways for the oxidation of 2-octanone with nitric acid.
R = CH3(CH2)4—
RCH2C-CH3
OH
H
RCH2C-CH3
O
+ HNO2
+ HNO3
- H2O
- HNO3
+ HNO3
-2HNO2
RCH2C-CH3
ONO
H
RCH2C-CH3
O
+ HNO2
+ HNO3
- H2O
- N2O
RC-OH +
O
CH3COOH
O
RCH2C-OH + HCOOH

Chapter 2
22
The nitric acid oxidation of 2-octanone is studied simultaneously with the
oxidation of 2-octanol. Van Asselt and van Krevelen [1963a] found different
products when oxidizing cyclohexanone with nitric acid and nitrite, compared to
the oxidation of cyclohexanol. This probably has been caused by side reactions
with the NO2 formed, when a large amount of nitrite is added. The oxidation of
2-octanone is accompanied by the formation of small amounts of unidentified
and unstable compounds. These compounds were too unstable to be isolated and
identified. The simplified reaction pathways can be represented as in Figure 2.
Depending on the carbon bond broken, hexanoic acid and acetic acid or
heptanoic acid and formic acid are formed. The amount of hexanoic acid as
found experimentally is approximately two times the amount of heptanoic acid.
The formic acid may further react to CO2, see Longstaff and Singer [1954].
During the reaction nitrous acid and nitric acid are consumed.
In the description of the oxidation reactions it is assumed that the reaction
proceeds only via the nitrosonium ion NO+
. However, at high temperatures
above 60 ºC, the oxidation is known to proceed via a radical mechanism, see
Castellan et al. [1991]. This is outside the operating conditions that will be
applied.
2.3 Derivation of overall conversion rates
The determination of unambiguous stoichiometry and kinetic parameters for
oxidation reactions is impossible due to the lacking knowledge of the exact
composition of the inorganic compounds in the aqueous reaction phase and the
unidentified and unstable intermediates in the organic phase. Hugo and Mauser
[1983] confirmed this for the nitric acid oxidation of acetaldehyde. Therefore, it
has been chosen to derive semi-empirical equations for the conversion rates and
heat production rates.
The oxidation of 2-octanol (A) to 2-octanone (P) and further oxidation products
(X) is simplified to the following two reactions:
A B P B rnol+ → + 2 (4)
P B X rnone+ → (5)
where B represents the nitrosonium ion which accounts for the autocatalytic
behavior. The reactions with the nitrosonium ion take place in the aqueous nitric

23
acid phase, so also the mass transfer rates of the organic compounds have to be
taken into account.
This process is schematically represented in Figure 3. The liquid-liquid system
consists of an aqueous acid phase (Aq) with nitric acid and the reacting
nitrosonium ion (B), and an organic phase (Org) containing mainly 2-octanol
(A), 2-octanone (P) and further oxidation products (X).
Figure 3: Schematic representation of mass transfer with chemical reaction
during the oxidation with nitric acid. Concentration profiles near the liquid-
liquid interface for a slow reaction and low solubility.
The 2-octanol (A) diffuses through the organic phase via the interface into the
aqueous acid phase. In the boundary layer and/or bulk of the aqueous phase it
reacts with the nitrosonium ion (B) to form 2-octanone (P). The 2-octanone may
react with the nitrosonium ion (B) to form carboxylic acids (X) or it is extracted
to the organic phase.
In case the transport of the organic compound in the reaction phase is not
chemically enhanced and the concentration drop over the film in the reaction
CB,Aq
CA,Org
*
JP
CA,Org
Organic
phaseInterface
film
Aqueous
phase
x = δ x = 0
JA
CP,Org
*
CP,Org
CA,Aq
*CA,Aq
CP,Aq
*
CP,Aq

Chapter 2
24
phase being relatively small, it is possible to derive an overall reaction rate
expression, see Steensma and Westerterp [1990]:
r k C Ci eff i Aq B Aq= −( )1 ε , , (6)
where ( )1− ε refers to the volume fraction of the aqueous reaction phase; keff is
the effective reaction rate constant. Equation (6) can be used under the following
conditions:
• The rate of chemical reaction is slow with respect to the rate of mass transfer,
the rate of mass transfer is not enhanced by reaction, and the reaction mainly
proceeds in the bulk of the reaction phase. One must check that the consumption
by reaction in the thin boundary layer is negligible, which is justified if Ha < 0 3.
holds, see Westerterp et al. [1987]. The Hatta number Ha is defined as:
Ha
k C D
k
eff B Aq i
L Aq
=
,
,
(7)
and Di is the diffusivity of the organic compound A and kL Aq, the mass transfer
coefficient for A, both in the aqueous phase.
• The solubility of the organic compound in the aqueous phase is so low, that
mass transfer limitations in the organic phase can be neglected. At the interface
holds C mCi Aq i Org.
*
.
*
= .
• The concentration drop over the film of the organic component transferred is
less than 5%, see Steensma and Westerterp [1990], so C Ci Aq i Aq,
*
,≈ can be
assumed.
If these conditions are fulfilled the conversion rate is independent of the
hydrodynamic conditions and interfacial area, hence independent of the stirring
rate. The conversion rates are determined by the kinetics of the homogeneous
chemical reactions, which can be described by the effective reaction rate
constants keff,nol and keff,none for the oxidations of 2-octanol and 2-octanone,
respectively.

25
Kinetic expressions
The effective reaction rate expressions should also account for the effect of
temperature and the acid concentration. Oxidation reactions with nitric acid
solutions are usually very sensitive towards the acid strength, see Ogata [1978].
The influence of the acid strength can be accounted for with the Hammett’s
acidity function, H0, see e.g. Rochester [1970]. So the kinetic constant becomes:
k T H k
E
RT
m Heff eff
eff
Ho eff( , ) exp, ,0 0= − −
%&'
()*∞ 2 7 (8)
For this expression the preexponential factor, k eff∞, , the energy of activation,
E Reff / , and Hammett’s coefficient, mHo eff, , have to be determined
experimentally.
Conversion rates in a semi-batch reactor
In a semi-batch operation, where 2-octanol is fed to a reactor initially loaded
with nitric acid, the overall balances list:
- for the 2-octanol, A:
dn
dt
C r VA
dos A dos nol r= −ϕ , (9)
where ϕdos is the volumetric flow rate of the feed dosed into the reactor.
- for the 2-octanone, P:
dn
dt
r V r VP
nol r none r= − (10)
- for the carboxylic acids, X:
dn
dt
r VX
none r= (11)
- for the nitrosonium ion, B:
dn
dt
r V r VB
nol r none r= − (12)
- for the nitric acid, N:
dn
dt
r V r VN
nol r none r= − − (13)

Chapter 2
26
The yields are defined, on the basis of the total amount of 2-octanol fed, nA1:
ζ P
P
A
n
n
=
1
ζ X
X
A
n
n
=
1
ζ B
B
A
n
n
=
1
The mass balances above can be made dimensionless, see Chapter 3 for the
derivation, as follows:
d
d
m k t C
d
d
P
A eff nol dos A dos P X
P B Xζ
θ
θ ζ ζ
ζ ζ
θ
ζ
θ
= − −
+
−, , 1 6 0
(14)
d
d
m k t CX
P eff none dos A dos P
P Bζ
θ
ζ
ζ ζ
θ
=
+
, , 1 6 0
(15)
in which θ is the dimensionless dosing time t/tdos. After the end of the dosing
θ=1 in Equations (14) and (15) and the reaction proceeds as in a batch reactor.
ζ B0 is the initial concentration of nitrosonium ion which will be formed after
addition of the initiator. The boundary conditions for these differential equations
and the corresponding heat balance will be discussed later.
It is assumed the volumes of the aqueous phase and the organic phase are not
affected by reaction. During the oxidation of 2-octanol and 2-octanone the
average molecular weight of the organic compounds does not change much, so
this assumption is justified. The assumption of low solubility of reactants and
products in the aqueous phase, which also may result in a change in volume, has
to be validated.
In the simplified representation of the oxidation reactions, Equations (4) and (5),
the reactions can be described with only two dimensionless partial mass
balances. The model of Equations (9)-(15) will be used to obtain the relevant
kinetic parameters and to simulate the experimental conversion rates.

27
2.4 Experimental set-up
Reaction calorimeter
The oxidation reactions have been studied in a reaction calorimeter RC1 of
Mettler Toledo, which contains a jacketed reactor vessel. Using the reaction
calorimeter the flow of the heat Qcool is determined, which is transferred through
the wall of the vessel and which is proportional to the temperature difference
between the reactor contents Tr and the coolant temperature Tcool :
Q UA T Tcool cool r= ⋅ −1 6 (16)
The proportionality factor UA has to be determined by calibration, which is done
by introducing via an electrical heating element a known amount of energy QC:
UA
Q
T T
C
r cool
=
−1 6 (17)
The reaction calorimeter enables an accurate measurement of the temperatures
of the reactor contents and of the coolant. The heat balance for the reactor
operating in the semi-batch mode can be written as:
dT
dt
dT
dt
Q Q Q Q Qr
r
w
w R dos cool stirΓ Γ+ = + + + + ∞ (18)
where Γr is the thermal capacity of the reaction mixture and internal devices in
the reactor, and Γw is the thermal capacity of the reactor wall. The wall
temperature is estimated by: T T Tw r cool= +1
21 6. The different heat flows taken
into account are QR by the chemical reaction, Qdos by mass addition, Qcool to the
coolant, Qstir by the agitation and Q∞ to the surroundings.

Chapter 2
28
Experimental set-up and experimental procedure
The experimental set-up is shown in Figure 4. The RC1 (1) contains a jacketed
1-liter glass vessel of the type SV01. The main dimensions of the reactor are
given in Figure 5. The reactor content is stirred by a propeller stirrer with a
diameter of 0.04 m. The stirring speed is adjusted to 700 rpm. For further details
and drawings of the RC1 see Reisen and Grob [1985] and Mettler-Toledo
[1993].
Figure 4: Simplified flowsheet of experimental set-up. Ti: temperature
indicator; FC: flow controller.
The reactor is operated in the semi-batch mode under isothermal conditions. To
operate below room temperature an external cryostatic bath (2) of the type
Haake KT40 has been installed. Before the experiment is started, the equipment
is flushed with N2. The reactor is initially filled with 0.4 kg of HNO3-solution.
First the effective heat transfer coefficient is determined with the electrical
heater with a thermal power of 5 W. After that a small amount of 0.1 g NaNO2 is
added as initiator. As soon as the temperature of the reactor has reached a
constant value, the feeding of reactant 2-octanol is started by activating the
dosing system. The dosing system contains the supply vessel, which is located
H2O
H2O
5
8
Ti
1
2
6
3
4
FCTi
7

29
Dbaffles = 0.1Dvessel
Dstirrer = 0.04 m
Dvessel, min = 0.06 m
hcone = 0.16 m
αcone = 18º
Dbaffles
Dstirrer
Dvessel, min
hcone
αcone
on a balance of the type Mettler pm3000 (3), a Verder gear pump (4) and a
Mettler dosing controller RD10 (6). The feed rate is kept constant in the range of
0.05 to 0.4 kg/h. The nitric acid and organic solutions are immiscible and form a
dispersion. The nitric acid remains the continuous phase during the whole
experiment. During the oxidation of 2-octanol NOX-gases are formed, which
accumulate above the reaction mixture and are let off through an opening in the
reactor lid to the scrubber (5) to be washed with water. After addition of 0.1 kg
2-octanol the dosing is automatically stopped and the experiment is continued
for at least two times the total dosing time. The experiment is then brought to an
end by heating up the reactor contents to complete the conversion and after that
again a determination of the effective heat transfer coefficient.
Also the temperatures of the feed and of the surroundings are measured and
together with the feed flow rate monitored and stored by a computer. When the
reactor temperature exceeds a certain value the computer automatically triggers
an emergency cooling program and opens the electric valve in the reactor
bottom to dump the reactor content and quench it in ice (8). During an
experiment 4 to 10 samples of the dispersion are taken via a syringe, as
indicated by (7) in Figure 4.
Figure 5: Dimensions of the SV01 glass reactor.

Chapter 2
30
During an experiment samples of the dispersion are taken of approximately 1
ml, using a syringe. The dispersion, once in the syringe, separates directly in two
phases. The total amount of strong and weak acids in the aqueous phase is
determined by titration with a 0.1 M NaOH-solution in an automatic titration
apparatus of the type Titrino 702 SM of Metrohm. During the reaction some
unstable and unidentified compounds are formed and the composition of an
untreated sample changes with time. Therefore, the samples of the organic phase
are contacted with demineralized water to stabilize the sample and remove the
nitric acid from the organic phase. The organic phase is then analyzed by gas
chromatography using a Varian 3400 with a FID detector. The injector and
detector temperatures are set at 240 ºC. The column is packed with Carbopack C
and is operated at 190 ºC with N2 as carrier gas. The concentrations of 2-octanol,
2-octanone, hexanoic acid and heptanoic acid are determined using reference
samples and an integrator of type HP3392A.
To study the influence of temperature the oxidation reaction has been
investigated in the temperature range of 0 ºC to 40 ºC, for dosing times of 900 to
7200 s, for 100 g of 2-octanol and an initial nitric acid concentration of 60 wt%.
Furthermore a series of experiments has been carried out in the range of 50 to 65
wt% with a dosing time of 1800 s to study the influence of the initial nitric acid
concentration. A total of 33 runs were carried out to obtain kinetic data.
An example of an experimental run is shown in Figure 6. Two peaks can be
observed in the temperature of the reactor as a function of time. The first peak is
small and is caused by the addition of the initiator. The second one is caused by
the start of the reaction; its deviation from the temperature set remains usually
below 2 ºC for a dosing times of 30 minutes and longer. Deviations from
isothermicity were larger for experiments with a short dosing time of 15
minutes. In this case, at temperatures above 25 ºC the heat production rate was
so large that isothermal operation became impossible. In Figure 6b the
calculated heat production rate is plotted as function of time. The maximum in
the heat production rate is an easily to be detected, sensitive measure of the
course of the reaction. It will be used in some comparisons further on.

31
0
5
10
15
20
25
-2000 0 2000 4000 6000 8000
time [s]
Temperature[ºC]
addition initiator
start dosing
Treactor
Tcooling
stop dosing
Treactor setpoint
-25
0
25
50
75
100
-2000 0 2000 4000 6000 8000
time [s]
HeatflowQR[W]
start dosing
stop dosing
Qmax
a.
b.
Figure 6: Example of an isothermal semi-batch experiment at 20 ºC with an
initial load of 0.4 kg 60 wt% HNO3 and 0.1 g NaNO2. Addition of 0.1 kg 2-
octanol in a dosing time of 30 min.
a. Measured temperature of reactor contents and cooling jacket
b. Measured heat flow

Chapter 2
32
For the same experiment the molar amounts of the organic compounds in the
organic phase and the total molar amounts of weak and strong acids in the
aqueous nitric acid solution are given as a function of time in Figure 7. 2-
Octanol accumulates in the reactor and a part of the dosed 2-octanol reacts to 2-
octanone, which is partly converted into carboxylic acids. As a result, the yield
of 2-octanone exhibits a maximum.
The distribution of 2-octanol and 2-octanone has been estimated on the basis of
TOC analysis of a saturated 60 wt% nitric acid solution and mA = 0.005 and mP
= 0.006 for 2-octanol and 2-octanone, respectively. The distribution coefficients
of the carboxylic acids are estimated on the basis of gas chromatography
analysis and m ≈ 0.01 for both heptanoic acid and hexanoic acid and m ≈ 1.5 for
acetic acid. Thus, in view of the low solubilities for 2-octanol, 2-octanone,
heptanoic acid and hexanoic acid, the amounts of organic compounds in the
aqueous phase can be neglected. The simultaneously formed acetic and formic
acids will be distributed over both the organic phase and aqueous phase and, as a
result, the volume of aqueous phase will increase as the reaction proceeds. At
the same time a considerable quantity of nitric acid will dissolve into the organic
phase. The overall effect on the volume ratio is small, since hardly any change
in volume is observed during the experiments.
The aqueous phase contains strong and weak acids. The strong acid is nitric
acid, the different weak acids could not be distinguished in the titration method
used. The weak acids probably consist of acetic and formic acids as well as an
amount of inorganic acids like HNO2.
Due to the extraction of nitric acid a part is not available for reaction. The
amount of nitric acid in the organic phase is determined by titration with a 0.1 M
NaOH solution and is approximately 2.5 mol/kg organic phase for 50 to 60 wt%
HNO3. Therefore the amount of strong acid in the aqueous phase, determined by
titration as shown in Figure 7b, appears to decrease faster then one may expect
based on the stoichiometry of the reactions.

33
0
0.2
0.4
0.6
0.8
0 2000 4000 6000 8000
time [s]
Numberofmoles
2-Octanol
Carboxylic acids
2-Octanone
0
1
2
3
4
0 2000 4000 6000 8000
time [s]
Numberofmoles
Strong acids (e.g. HNO3)
Weak acids (e.g. HNO2, organic acids)
a.
b.
Figure 7: Molar amount as function of time for same run as in Figure 6.
a. Organic compounds in the organic phase;
b. Weak and strong acids in the aqueous nitric acid phase.

Chapter 2
34
2.5 Experimental results
The kinetic parameters of the proposed model can be found by measuring the
conversion rates by means of thermokinetic measurements in the calorimeter in
combination with chemical analyses. Before the kinetic parameters are evaluated
the reaction regime has to be identified.
Identification of reaction regime
Effect of agitation
If the conversion rate in a liquid-liquid reaction is not influenced at all by mass
transfer resistances, it should be independent of the interfacial area and, hence,
of the degree of agitation. The influence of the stirring rate on the conversion
rate has been experimentally determined at 20, 30 and 40 ºC.
In Figure 8 the measured maximum heat production rate is plotted against the
stirring speed. The maximum heat production initially increases with stirring
speed, but becomes independent of the agitation above 300 rpm. At a stirring
speed below 150 rpm the reaction mixture separates into two liquid phases and it
becomes well dispersed at stirring rates above 500 rpm, as can be visually
observed. Between 150 and 500 rpm a certain volume of undispersed organic
phase is visible above the dispersion and the heat production rates fluctuate in
time. For a stirring rate of above 500 rpm evidently the mass transfer resistance
1/kLa does not play a role anymore. Therefore, a stirring rate of 700 rpm has
been chosen for all experiments.
Effect of phase volume ratio
By assuming the nitrosonium ion being the reactive species it is likely that the
reaction takes only place in the aqueous acid phase. The conversion rate is
usually proportional to the volume of reacting phase, according to: R kC C VA B R= ,
where CA and CB are the concentrations of the reacting compounds in the
reaction phase with volume VR. On the other hand, the reaction phase can be
identified by varying the volume of the phases and keeping all other parameters
constant, see e.g. Atherton [1993] and Hanson [1971].

35
0
50
100
150
200
0 200 400 600 800 1000 1200
Stirring speed [rpm]
Maximumheatproductionrate[W]
40ºC
30ºC
20ºC
Figure 8: Maximum heat production rate versus stirring speed at 20, 30 and 40
ºC. Isothermal semi-batch experiments with an initial load of 0.4 kg 60 wt%
HNO3 and 0.1 g NaNO2. Addition of 100 g 2-octanol in a dosing time of 30 min.
However, for the autocatalytic reaction, complications arise when the
concentration of nitrosonium ion CB has to be kept constant, while the volume of
the aqueous phase VR is changed. The number of moles of nitrosonium ion nB =
CBVR is equal to the number of moles of product in the non-reaction phase nP =
CPVd. The concentration of nitrosonium ion is therefore equal to CB = CPVd/VR
and consequently the conversion rate is also equal to R = kCACPVd. Thus a larger
initial volume of aqueous phase VR will be accompanied by a lower
concentration of nitrosonium ion CB and as a result there is no change in
conversion rate.

Chapter 2
36
Run Volume of acid
phase
[ml]
Volume of
organic phase
[ml]
Feed concentration
2-octanol
[mol/l]
1 293 120 6.40
2 450 120 6.40
3 525 120 6.40
4 295 150 4.98
5 295 173 4.33
6 295 225 3.64
7 295 278 2.77
Table 1: Experimental conditions of isothermal experiments with
varying concentration and volumes. All experiments with
initially 60 wt% HNO3 and 0.1 g NaNO2 at 25 ºC, in the semi-
batch mode with a dosing time of 30 minutes.
The oxidation reaction has been carried out with different volumes of the
aqueous reaction phase as is shown in Table 1. The experimental results are
plotted in Figure 9 and show an increase in heat production rate with an
increasing volume of nitric acid. This increase in the maximum heat production
rate can be explained entirely by the effect of the acid strength on the kinetic
constant k: the nitric acid remains at a higher concentration level for a larger
initial volume, as its excess is larger. Thus a larger volume of reaction phase VR
has no effect on the part CACBVR as mentioned above. This confirms nitric acid
being the reaction phase.
This can be double-checked by changing the volume of the organic phase, which
can be increased by diluting the 2-octanol with inert hexane, keeping the total
amount of 2-octanol constant. The results of these experiments are shown in
Figure 10. The maximum conversion rate decreases, when the amount of
organic non-reacting phase is increased. This can be explained, partly by the
lower concentration of the 2-octanol and 2-octanone in the aqueous phase and
partly, by a lower concentration of the nitrosonium ion, as also mentioned by
Ogata et al. [1967].
The above phenomena also support the assumed ionic mechanism via NO+
in
the aqueous acid phase. Thus, although some reaction may take place in the
organic phase its contribution to the overall rate will be neglected. So it is
assumed that the reaction only takes place in the aqueous, nitric acid phase.

37
60
80
100
120
0.1 0.3 0.5 0.7
Volume of aqueous phase [l]
Qmax.[W]
20
40
60
80
100
0.1 0.15 0.2 0.25 0.3
Volume organic phase [l]
Qmax.[W]
Figure 9: Maximum heat production rate versus volume aqueous nitric acid
phase. Isothermal semi-batch experiments with an initial load of 60 wt% HNO3
and 0.1 g NaNO2. Addition of 0.1 kg 2-octanol in a dosing time of 30 min.
Figure 10: Maximum heat production rate versus volume organic phase.
Isothermal semi-batch experiments with an initial load of 0.4 kg 60 wt% HNO3
and 0.1 g NaNO2. Addition of 2-octanol in hexane as indicated in Table 1.

Chapter 2
38
Determination of kinetic parameters
Now the kinetic parameters can be determined using the conversion rate
expressions for slow liquid-liquid reactions, provided the heats of reaction are
known.
Determination of effective heats of reaction
The heat production is determined by the chemical reactions and physical
phenomena like dilution, etc. The heat production rate by n chemical reactions
can be written as:
Q r H VR i i
i
n
r= ∑ ∆ (19)
The amount of heat released by the reaction ∆Ε is determined by integrating the
experimentally measured heat generation rate QR over the reaction time:
∆E Q dt Q Q dtcalorimeter R
t
nol none
t
= = +I I0 0
1 6 (20)
where Qnol and Qnone are the heat generated by the oxidation of 2-octanol and 2-
octanone, respectively. The results of the chemical analyses are used to calculate
the amounts of heat generated by both reactions separately:
∆ ∆ ∆E H n H nanalyses eff nol P X A eff none X A= ⋅ + ⋅ + ⋅ ⋅, ,ζ ζ ζ1 6 1 1 (21)
The effective heats of reaction ∆Heff,nol and ∆Heff,none are obtained using the
complete set of isothermal experiments and by minimizing the deviation
between the amount of heat measured by the calorimeter, ∆Εcalorimeter, and the
amount of heat calculated using the yields, ∆Εanalyses. The results are listed in
Table 2.
Reaction ∆Heff
[kJ/mol]
∆Hcalc
[kJ/mol]
2-octanol Æ 2-octanone, ∆Heff,nol 160 150
2-octanone Æ products, ∆Heff,none 520 620
Table 2: Experimentally determined effective heats of reaction
∆Heff and calculated ∆Hcalc based on the heats of formation.

39
0
100
200
300
400
500
0 1800 3600 5400 7200
Time [s]
∆Ε[kJ]
∆Ηeff, none/∆Ηeff, nol = H = 3.25
∆Εnone
∆Εnol
1.1•H
0.9•H
∆Εnol+∆Enone
∆Εcalorimeter
Figure 11: Amount of heat generated as a function of time by the oxidation of
2-octanol ∆Enol and 2-octanone ∆Enone as measured in the calorimeter, and as
calculated on the basis of the concentration time profiles.
The heat generated as a function of time is shown for a single run in Figure 11,
where the heat generated by the separate reactions ∆Enol and ∆Enone and the total
amount of heat generated ∆Eanalyses = ∆Enol + ∆Enone using Eq.(21) or ∆Ecalorimeter
using Eq.(20), respectively, are displayed. The ratio of the effective heats of
reaction, H H Heff none eff nol= ∆ ∆, ,/ , is equal to H = 3.25. In the same figure are
shown the calculated amount of heat ∆Ε with 0.9H and 1.1H respectively. For
this single run the amount of heat ∆Eanalyses calculated with the conversions is in
agreement with ∆Ecalorimeter measured by the calorimeter, during the time of the
experimental run.
A comparison between the calculated heat production and the experimental
determined heat production for all runs is given in Figure 12. Although the
points do not seem completely random by distribution, the deviations are small
and the values of ∆Heff,nol and ∆Heff,none are acceptable.

Chapter 2
40
10
100
1000
10 100 1000
Amount of heat ∆Qanalyses [kJ]
Amountofheat∆Qcalorimeter[kJ]
Figure 12: Parity plot of calculated amount of heat generated according to
Eq.(21) and in the calorimeter experimentally determined amount of heat
produced, Eq.(20), for all runs.
An approximate estimate of the heats of reaction can be made using the heats of
formation of the reacting species as depicted in Figure 1 and Figure 2. For the
oxidation of 2-octanol to 2-octanone the calculated heat of formation is in good
agreement with the experimentally determined reaction heat. For the oxidation
of 2-octanone to carboxylic acids a 16% difference was found; this is probably
the result of endothermic decomposition reactions, which produce NOX-gases,
and which have not been taken into account.
Determination of the model parameters
The kinetic constants for the proposed model can now be found by comparing
the experimental conversion rates of 2-octanol and 2-octanone and the proposed
model equations. During an experiment the conversion rates can be determined
by evaluating the heat flow measurements or the results of the chemical

41
0
0.2
0.4
0.6
0.8
1
0 0.5 1 1.5 2
θ = t/tdos [-]
Concentration[-]
nX
nA1
nP
nA1
analyses, using Equation (19) and the determined effective heats of reaction as
listed in Table 2. The total heat production rate in the reactor QR is equal to:
Q Q Q r V H r V HR nol none nol r eff nol none r eff none= + = ⋅ + ⋅∆ ∆, , (22)
On the basis of the chemical analyses the conversion rates can be obtained by
differentiation of a polynomial fit of the measured data points, as is shown in
Figure 13 and using the following equations:
r V
n
t
d
d
d
d
nol r
A
dos
P X
1 6= +

ζ
θ
ζ
θ
and (23)
r V
n
t
d
d
none r
A
dos
X
1 6=
ζ
θ
(24)
Figure 13: Measured concentrations by chemical analysis (dots) and polynomial
function (lines) for a single run.
The sampling frequency during an experiment was usually once per 15 minutes,
which results in 5 to 10 samples per run. Due to this limited amount of sampling
data points, not always a useful polynomial expression could be obtained for the

Chapter 2
42
2-octanone (P) concentration. The concentration of the further oxidation
products (X) increases approximately linearly with time under the experimental
conditions applied and good polynomial functions could be found, as shown in
Figure 13. To improve upon the accuracy of the conversion rate of 2-octanol
rnolVr the total conversion rate from the heat flow measurements QR is combined
with the information of chemical composition of the further oxidation products
(X) as function of time. The conversion rate of 2-octanol rnolVr can also be
expressed as:
r V
Q r V H
Hnol r
R none r eff none
eff nol
1 6 2 7=
− ⋅∆
∆
,
,
(25)
For every run in the reaction calorimeter first the conversion rate of 2-octanone
r Vnone r is evaluated using Equation (24) and the polynomial expression. Then the
conversion rate of 2-octanol r Vnol r is evaluated by Equation (25).
The conversion rates can also be found after combining the conversion rates
from Equation (23) and (24) with the mass balances Equation (14) and (15):
r V
n
t
m k t Cnol r
A
dos
A eff nol dos A dos P X
P B
1 6 1 6= − −
+

, , θ ζ ζ
ζ ζ
θ
0
(26)
r V
n
t
m k t Cnone r
A
dos
P eff none dos A dos P
P B
1 6 1 6=
+
, , ζ
ζ ζ
θ
0
(27)
All parameters in the Equations (26) and (27) are known, except mAkeff,nol and
mPkeff,none. The kinetic constants of the proposed expression of Equation (8) are
obtained by non-linear regression using the complete set of isothermal
experiments and fitting the Equations (26) and (27) to the results of Equations
(24) and (25). The results determined in the range of 0 to 60 ºC and acid strength
of H0 = 2.4 to 3.5 are listed in Table 3. The standard deviation of the
experimentally determined reaction rate constants compared to the calculated
ones is 60%. The accuracy will be visualized in the following.
Reaction mkW,eff
[l/mol s]
Eeff/R
[K]
mHo,eff
[-]
2-octanol Æ2-octanone 1 · 105
11300 6.6
2-octanone Æ products 1 · 1010
12000 2.2
Table 3: The effective reaction rate constants for the
oxidation reactions.

43
1.E-07
1.E-06
1.E-05
1.E-04
1.E-03
1.E-02
1.E-01
1.E+00
2.1 2.6 3.1 3.6
-H0 [-]
mkeff[m
3
/kmols]
2-octanol 2-octanone
2-octanone carboxylic acids
10
0
10
-1
10
-2
10
-3
10
-4
10
-5
10
-6
10
-7
The effective kinetic constant depends on temperature and acid strength. To
discuss the influence of these parameters on the kinetic constants the value of
mkeff is measured for both reactions. The kinetic constant is very sensitive to the
nitric acid concentration: below 40 wt% the reaction is so slow that hardly any
heat production is measurable, while above 65 wt% the reaction becomes too
fast. Expressed as an exponential order in the concentration of HNO3, the
exponent would be as high as 12 for the oxidation of 2-octanol. This has no
physical or chemical meaning, so Hammett’s acidity function is used, see
Rochester [1970]. Figure 14 shows a plot of mkeff at 20 ºC as a function of
Hammett’s acidity function H0. The slope of ln(mkeff) versus -H0 is 1.25 and 0.41
for the oxidations of 2-octanol and 2-octanone, respectively. These values can
be compared to those reported in literature. Ogata et al. [1966] found a slope of
0.95 for the nitric acid oxidation of benzyl alcohol, while for the oxidation of
benzaldehyde a value of 0.43 has been reported, see Ogata et al. [1967]. The
oxidation of 2-octanol depends more strongly on the nitric acid concentration
then the oxidation of 2-octanone. This has also been found for the oxidation of
benzyl alcohol and benzaldehyde respectively as described above. Therefore, to
increase the yield of 2-octanone the concentration of nitric acid should be high.
The term mHo,eff accounts for the acidity effect on the conversion rate including
the acidity influence on the solubility, which is known to increase with
increasing HNO3 concentration, see Rudakov et al. [1994].
Figure 14: Effect of acid strength on the reaction rate constants for the
oxidation of 2-octanol and 2-octanone, respectively. Lines calculated according
to Eq.(8) and parameters from Table 3 for T = 20 ºC.

Chapter 2
44
1.E-07
1.E-06
1.E-05
1.E-04
1.E-03
1.E-02
1.E-01
1.E+00
1.E+01
2.8 3.0 3.2 3.4 3.6 3.8
1000/T [1/K]
mkeff[m
3
/kmols]
2-octanol 2-octanone
2-octanone carboxylic acids
10
0
10
-1
10
-2
10
-3
10
-4
10
-5
10
-6
10
-7
10
1
In Figure 15 the value of mkeff is plotted at 60 wt% HNO3 as a function of
temperature. The term Eeff/R accounts for the temperature influence on the
conversion rate, including the temperature influence on the solubility and, more
important, the Hammett acidity. The latter is only well tabulated for HNO3-
solutions at 25 ºC, see Rochester [1970], but some data points at 20 ºC indicate
an increasing acidity with increasing temperature, hence the value of Eeff/R is
overestimated.
Although no experimental data on the oxidation of 2-octanol or 2-octanone have
been published, comparable data can be found in literature for other nitric acid
oxidations. The reported data on energy of activation vary from 9000 K for the
oxidation of methoxyethanol, see Strojny [1971], to 14230 K for benzyl alcohol,
see Ogata et al. [1966]. The same range is found for aldehydes or ketones: from
8000 K for cyclohexanone, see van Asselt and van Krevelen [1963c] to 14400 K
for benzaldehyde, see Ogata et al [1967]. When the determined values of mkeff
for both reactions are compared, an equal trend is observed with respect to
temperature. As the energy of activation has comparable values for the oxidation
of alcohols, aldehydes or ketones, selectivity can not be influenced by
temperature.
Figure 15: Effect of temperature on the reaction rate constants for the
oxidation of 2-octanol and 2-octanone, respectively. Lines calculated according
to Eq.(8) and parameters from Table 3 for 60 wt% HNO3.

45
2.6 Simulation of isothermal runs
The mathematical model for the oxidation rates has been tested using the kinetic
parameters as described above. The mass balances Equation (14) and (15) are
expressed as two differential equations and can be solved simultaneously using a
fifth order Runge-Kutta method with an adaptive step size control, see Press et
al. [1986]. In view of the autocatalytic behavior, whereby some reaction product
must be present before the reaction can start, an initiator has to be added. For all
experiments an addition of 0.1 g NaNO2 has been chosen. This is, as
experimentally found, the minimum amount to be added to ensure the reaction
starts immediately. To solve the differential equations and to account for the
initial reaction rate, an initial concentration of nitrosonium ion ζB0 has to be
taken, which is an optimizing problem. The initial reaction rates as
experimentally determined and calculated are in good agreement provided an
initial concentration of nitrosonium ion equal to 3.5% is taken. Thus, the
boundary conditions for these differential equations are: ζP0 = 0, ζX0 = 0 and ζB0
= 0.035 at θ = 0. The differential equations together with the kinetic parameters
in Table 3 can now be used to simulate the experiments.
Figure 16 shows the experimentally determined and simulated heat production
rates as a function of time. The simulated heat production rates Qnol and Qnone are
plotted for the separate reactions. Also both, the simulated and experimental,
total heat production rates Q Q QR nol none= + are plotted. The measured and
simulated conversion-time profiles for 2-octanol, 2-octanone and carboxylic
acids are shown in Figure 17 for the same series. The 2-octanol was added in 30
minutes to 60 wt% HNO3 at a temperature of 10, 20 and 40 ºC respectively. One
can observe that the heat generation rate increases with increasing temperature,
which is the result of both the increasing conversion rate of 2-octanol as well as
the increasing rate of the more exothermic oxidation of 2-octanone.

Chapter 2
46
0
50
100
150
200
0 1800 3600 5400 7200
Time [s]
Heatproductionrate,Q[W]
Qnol + Qnone
Qnone
Qnol
QR, experimental
0
25
50
75
100
0 1800 3600 5400 7200
Time [s]
Qnone
Qnol
QR, experimental
Qnol + Qnone
0
25
50
75
100
0 1800 3600 5400 7200
Time [s]
Qnone
Qnol
QR, experimental
Qnol + Qnone
Figure 16: Experimental total heat production rate QR,experimental (thick line) and
simulated (thin lines) heat production rates Qnol, Qnone and QR,simulated= Qnol+Qnone.
Isothermal semi-batch experiments at a temperature of 10, 20 and 40 ºC
respectively, with an initial load of 0.4 kg 60 wt% HNO3 and 0.1 g NaNO2.
Addition of 0.1 kg 2-octanol in a dosing time of 30 min.

47
0
0.2
0.4
0.6
0.8
1
0 1800 3600 5400 7200
Time [s]
Numberofmoles
0
0.2
0.4
0.6
0.8
1
0 1800 3600 5400 7200
Time [s]
Numberofmoles
0
0.2
0.4
0.6
0.8
1
0 1800 3600 5400 7200
Time [s]
Numberofmoles
Figure 17: Experimental (dots) and simulated (lines) conversions of 2-octanol
(q, ), 2-octanone (s, ) and carboxylic acids (v, ). Isothermal semi-
batch experiments with experimental conditions as for Figure 16.

Chapter 2
48
0
0.2
0.4
0.6
0.8
1
0 1 2 3 4
Dimensionless time θ = t/tdos [-]
Concentration2-octanone[-]
20 ºC
40 ºC
60 ºC
nP
nA1
( )max
Figure 18: Concentration of 2-octanone as a function of time for isothermal
semi-batch experiments and the maximum concentration of 2-octanone as
obtained during each run. Simulations with a temperature of 20, 40 and 60 ºC
and further conditions as for Figure 16.
The conversion of 2-octanol increases with increasing temperature and as a
result the location of the maximum concentration of 2-octanone in the
conversion-time profile shifts towards shorter reaction times. The concentration
profiles of 2-octanone for simulations of isothermal runs at 20, 40 and 60 ºC are
plotted in Figure 18. In the same figure, the line is plotted connecting all the
maximum concentrations of 2-octanone. The maximum concentration of 2-
octanone is found after a long reaction time when the reactor temperature is low.
The energy of activation has comparable values for both reactions. Therefore,
the maximum concentration is hardly affected by the reactor temperature and
will be practically constant as long as the reaction time is sufficiently long.
At higher temperatures the location of the maximum concentration of 2-
octanone shifts towards shorter reaction times. The influence of dosing becomes
visible when the maximum concentration is obtained just after the dosing has
been stopped at θ = 1. In that case the maximum concentration decreases.
A comparison between simulations and experimental results shows the proposed
model is sufficiently accurate to describe the conversion and heat production

49
rates of the oxidation reactions. Especially, when one takes into account the
complexity of the oxidations reaction and the simplicity of the model.
2.7 Model validation and limitations
The process of mass transfer with chemical reaction during the oxidations of 2-
octanol and 2-octanone with nitric acid has been modeled by assuming that the
conversion rate is not affected by mass transfer rates. The verification of the
assumptions described in Section 2.3 regarding these mass transfer rates is
discussed below:
Slow reaction, Ha0.3
The Hatta numbers are calculated for both reactions and listed in Table 4 as a
function of temperature. These values have been obtained for CNaNO2 0, is 4.9·10-3
M, CHNO3 0, is 13.0 M and the stirring rate is 700 rpm. The diffusivity coefficients
have been calculated using the relation of Wilke and Chang [1955] together with
the relation of Cox and Strachan [1972] to correct for nitric acid mixtures. The
estimation of the mass transfer coefficients will be discussed in the next
paragraph.
Temperature
[ºC]
Hanol, max. Hanone, max.
0 0.2 0.02
10 0.3 0.02
20 0.4 0.06
30 0.5 0.07
40 0.6 0.09
Table 4: Calculated maximum Hatta numbers,
Hamax, for the isothermal oxidation experiments with
N = 700 rpm. Initial: 60 wt% HNO3, 0.1 g NaNO2.
The calculated Hatta numbers for the oxidation of 2-octanol to 2-octanone
indicate that the transfer rates are not enhanced by chemical reaction as long as
the temperature is below 20 ºC. The conversion rate of 2-octanone to further
oxidation products is not chemically enhanced in the whole range of applied
temperatures. If the reaction is not slow compared to mass transfer, the

Chapter 2
50
enhancement can be estimated by the expression of Danckwerts, see e.g.
Westerterp et al. [1987]:
E HaA = +1 2
(28)
The deviations are within 5% and 10% up to a temperature of 10 ºC and 20 ºC
respectively. The deviation is slightly higher at 40 ºC: 17%, but still reasonably
small as also experimentally demonstrated by the influence of stirring speed.
Mass transfer resistance in the organic phase negligible
The mass transfer resistance in the organic phase is zero if the phase consists of
pure reactant without solvent as in the case of the oxidation of 2-octanol. As the
reaction proceeds, 2-octanone is formed and dilutes the organic phase. Thus the
validity of the neglect of the mass transfer resistance in the organic phase must
be examined. This assumption holds, see Westerterp [1987], if:
k
k m
L Org
L Aq
,
,
1 (29)
The mass transfer coefficients kL,Aq for 2-octanol and 2-octanone in the
continuous, aqueous phase can be estimated with the empirical correlation of
Calderbank and Moo-Young [1961] as discussed in detail in Chapter 4. A
typical value of the mass transfer coefficients for both 2-octanol and 2-octanone
in the continuous phase is kL,Aq = 20·10-6
m/s for the range of experimental
conditions. This value is in agreement with the value reported by Chapman et al.
[1974]. They found experimentally kL = 10.3·10-6
m/s for toluene in a
HNO3/H2SO4 solution.
In view of the low solubility of the organic compounds in nitric acid with mA =
0.005 and mP = 0.006 for 2-octanol and 2-octanone, respectively, and the mass
transfer coefficient in liquid-liquid dispersions of the same order of magnitude,
see e.g. Laddha and Degaleesan [1976] and Heertjes and Nie [1971], this gives
for k k mL Org L Aq, ,( ) a value of approximately 200. Therefore, the mass transfer
resistance in the organic phase is negligible for the transport of both 2-octanol
and 2-octanone.
The concentration drop over the film is negligible
The concentration drop from Ci Aq,
*
to Ci Aq, is relatively more important if mass
transfer resistance in the aqueous phase is higher. When the concentration drop
is more than say 5%, the simple approximation C Ci Aq i Aq,
*
,≈ starts to lead to

51
inaccuracies, see Steensma and Westerterp [1990]. To check this approximation
it is possible to compare the rate of mass transfer with the chemical reaction, see
Zaldivar et al. [1995]:
Ja k C C aL i Aq i Aq= −,
*
,2 7 (30)
Ja k C Ceff i Aq B Aq= −( )1 ε , , (31)
where a is the interfacial area per unit volume of reactor content. The
combination of both equations gives:
1
1
−( )
= −
ε k C
k a
C
C
eff B Aq
L
i Aq
i Aq
, ,
*
,
(32)
Hence, in the case where C Ci Aq i Aq,
*
,≈ it must be checked whether
( ) ,1 1− ε k C k aeff B Aq L . The total interfacial area is estimated by means of the
Sauter mean drop diameter, d32, which is defined as:
d a32 6= ε / (33)
where ε is volume fraction of dispersed phase and a the interfacial area per unit
volume of reactor content. The average drop size depends upon the conditions of
agitation and the physical properties of the liquids. For baffled stirred tank
reactors the Sauter mean drop diameter d32 can be estimated using the
correlation:
d
D
A B We
stir
32 0 6
1= + −
( ) .
ε (34)
where Dstir is the impeller diameter, ε is the volume fraction of dispersed phase,
A and B are empirical constants, which must be determined experimentally for a
given reactor set-up and liquid-liquid system, see Chapter 4. We is the Weber
number, defined as:
We
N Dstir c
=
2 3
ρ
σ
(35)
where N is the stirring rate, ρc is the density of the continuous phase and σ is
the interfacial tension. Equation (34) has been used by numerous workers,

Chapter 2
52
whereby the values of A and B depend on the geometry. With the used values for
A and B reasonable values have been obtained for the drop size. This is
sufficiently accurate to estimate the validity of the concentration drop over the
film.
The interfacial tension is predicted using the empirical correlation of Good and
Elbing [1970]:
σ γ γ φ γ γ12 1 2 12 1 22= + − (36)
where φ12 is an experimentally determined interaction parameter and γ 1 and γ 2
are the surface tensions of the pure components. The interaction parameter φ12 is
not known for 2-octanol. Therefore the value for n-octanol has been used, see
Good and Elbing [1970], which is equal to φ12=0.97. The surface tensions for
both 2-octanol and 2-octanone are equal to 0.026 N/m at 20 ºC, see Daubert et
al. [1989], and for a 60 wt% HNO3 solution it is equal to 0.063 N/m, see
Zaldivar et al. [1996]. The liquid-liquid interfacial tension between 2-octanol, 2-
octanone or a mixture of both with a 60 wt% nitric acid solution is thus equal to
σ = 0.010 N/m. This can be compared to the experimental value between
octanol and water of σ = 0.0085 N/m, as measured by van Heuven and Beek
[1971].
Temperature
[ºC]
( ) , ,1− ε k C k aeff nol B Aq L ( ) , ,1− ε k C k aeff none B Aq L
0 0.02 0.0001
10 0.05 0.0004
20 0.07 0.001
30 0.15 0.004
40 0.20 0.006
Table 5: Validity of the assumption of a negligible
concentration drop over film for 2-octanol (reaction ‘nol’)
and 2-octanone (reaction ‘none’), respectively. Isothermal
oxidation experiments with N = 700 rpm and initially
60wt% HNO3 and 0.1 g NaNO2.
The Weber-number is now equal to We =1175. The interfacial area increases
with the hold-up of the organic phase for the used system from 8000 to 15000
m2
/m3
. Typical values of ( ) ,1− ε k C k aeff B Aq L are listed in Table 5 as a function of

53
temperature. The assumption of a negligible concentration drop over the film for
2-octanone is valid. For 2-octanol this is not true and the simple approximation
C Ci Aq i Aq,
*
,≈ leads to inaccuracies. The deviations are within 5% and 10% up to a
temperature of 10 ºC and 20 ºC respectively.
As can be concluded from Table 4 and Table 5, all assumptions are valid with
deviations below 10% as long as temperature is lower than 20 ºC. At a higher
temperature the description of the oxidation of 2-octanol using the reaction rate
expression of Equation (6) may lead to deviations of up to 20% at 40 ºC.
Fortunately, the deviations are small and still within the experimental error.
Thus the model based on the slow liquid-liquid reaction regime can be used
without introducing larger inaccuracies.
Model verification with isoperibolic experiments
The data from the isothermal experiments, being the concentrations versus time
and heat production rate versus time, were used to fit the reaction rate equations.
Data from isoperibolic experiments can be used to test the accuracy of the
derived kinetic expressions. The data from experiments with a constant jacket
temperature have not been used to determine the kinetic expressions.
The mathematical model with the mass balances Equation (14) and (15) together
with the heat balance Equation (18) now can be used to describe the temperature
profile. The isoperibolic experiments were carried out in the same way as the
isothermal runs, except that the calorimeter now is operated with a constant
jacket temperature. In Figure 19 the temperature profiles are plotted for five
isoperibolic experiments with different jacket temperatures: the experimental
profiles are in good agreement with the simulations. In Figure 20 the
temperature profiles are plotted for four isoperibolic experiments with different
jacket temperatures and a faster dosing rate. As can be seen one is working in a
parametric sensitivity region, where the maximum reactor temperature, Tmax, is
sensitive towards the cooling temperature Tcool.
Under these conditions even a small deviation between model and actual
parameters will lead to large discrepancies. At higher temperatures the model
overestimates the reactor temperature, which can be attributed to evaporation of
the nitric acid solution, which has not been incorporated in the model. However,
the simulated and the experimental results show the same thermal behavior. This
thermal behavior of the oxidation reaction will be discussed in more detail and
under varying experimental conditions in Chapter 3.

Chapter 2
54
0
20
40
60
80
100
120
0 0.5 1 1.5 2
theta [-]
Temperature[ºC]
0
10
20
30
40
50
60
0 0.5 1 1.5 2
theta [-]
Temperature[ºC]
Figure 19: Experimental (continuous line) and simulated (dotted lines) reactor
temperatures in some isoperibolic semi-batch experiments with varying coolant
temperature with T0 = Tcool. Initial load of 60 wt% HNO3 and 0.1 g NaNO2.
Addition of 100 g 2-octanol in a dosing time of 120 min.
Figure 20: Experimental (continuous line) and simulated (dotted lines) reactor
temperatures in some isoperibolic semi-batch experiments with varying coolant
temperature with T0 = Tcool. Initial load of 60 wt% HNO3 and 0.1 g NaNO2.
Addition of 100 g 2-octanol in a dosing time of 30 min.

55
2.8 Discussion and conclusions
The main objective of this chapter is to determine the kinetic parameters of the
model proposed to describe the heterogeneous oxidation of 2-octanol to 2-
octanone and the unwanted, further oxidation reactions to carboxylic acids. The
oxidation of 2-octanol and 2-octanone with nitric acid exhibits the typical
features of nitric acid oxidation reactions, like a long induction time without
initiator; autocatalytic reaction; strong dependence of mineral acid concentration
and high energy of activation, see Ogata [1978]. Although the main phenomena
of nitric acid oxidation reactions are well known the exact mechanism is still not
elucidated. There is a limited knowledge of the exact chemical structure of the
compounds in the aqueous reaction phase and of a number of unknown, unstable
compounds in the organic phase. As a consequence of this a strong model
reduction was necessary to describe the overall reaction rates. The model
reduction in this case gave satisfactory results, as also demonstrated by Hugo
and Mauser [1983].
The observed conversion rates of the complex reactions of 2-octanol and 2-
octanone with nitric acids can be correlated using only two kinetic equations, in
which the effect on temperature is described through the Arrhenius equation and
the effect on acid strength through Hammett’s acidity function. The
experimental results and simulations are in good agreement, hence the employed
film model is satisfactory.
The oxidation reactions have been studied in the range 0 to 40 ºC, with initial
nitric acid concentrations of 50 to 65 wt% and a stirring rate of 700 rpm. The
results indicate the oxidation of 2-octanol is operated in the non-enhanced
regime when nitric acid is below 60 wt% or when the temperature is below 25
ºC at 60 wt% HNO3, while the oxidation of 2-octanone is operated in the non-
enhanced regime for the whole range of experimental conditions considered.
Under these conditions the mass transfer resistance does not influence the
overall conversion rate, so the governing parameters are the reaction rate
constant and the solubility of the organic compounds in the nitric acid solution.
This has also been experimentally confirmed by determining the influence on
stirring rate.
Even though the kinetic constants have been determined only up to a
temperature of 40 ºC, the simulated results for isoperibolic experiments at higher
temperatures are still acceptable. Therefore it can be concluded that it has been
possible to describe the thermal behavior of the semi-batch reactor for the nitric
acids oxidation reactions with the film model for slow liquid-liquid reactions

Chapter 2
56
and a simplified reaction scheme. In Chapter 3 the thermal behavior of this
consecutive heterogeneous liquid-liquid reaction system will be further
evaluated.
Acknowledgements
The author wishes to thank S.E.M. Geuting, R.H. Berends, V.B. Motta, E.A.H.
Ordelmans and S.P.W.M. Lemm for their contribution to the experimental work,
and F. ter Borg, G.J.M. Monnink and A.H. Pleiter for technical support. W.
Lengton and A. Hovestad are acknowledged for the assistance in the analysis.
Notation
a Interfacial area per volume of reactor content = 6 32ε / d [m2
/m3
]
A Effective cooling area [m2
]
C Concentration [kmol/m3
]
CP Specific heat capacity [J/Kg K]
D Diameter [m]
DI Diffusivity coefficient component i [m2
/s]
d32 Sauter mean drop diameter [m]
EA Enhancement factor [-]
EAct Energy of activation [J/kmol]
h Height [m]
H ∆ ∆H Heff none eff nol, ,/ [-]
H0 Hammett’s acidity function [-]
Ha Hatta number [-]
J Mole flux [kmol/m2
·s]
kLaq Mass transfer coefficient in the aqueous phase [m/s]
kLorg Mass transfer coefficient in the organic phase [m/s]
keff Effective second order reaction rate constant [m3
/kmol·s]
k∞,eff Effective preexponential constant [m3
/kmol·s]
M Molecular weight [kg/kmol]
m Molar distribution coefficient [-]
mHo Hammett’s coefficient [-]
n Number of moles in the reactor [kmol]
N Stirring rate [s-1
]
Q Heat flow [W]
R Gasconstant = 8315 [J/kmol·K]
r Rate of reaction per volume of reactor content [kmol/m3
s]

57
t Time [s]
tdos Dosing time [s]
T Temperature [K]
U Overall heat transfer coefficient [W/m2
K]
V Volume [m3
]
Greek symbols
α Angle of cone [º]
∆H Heat of reaction [kJ/mol]
∆E Amount of heat [kJ]
ε Volume fraction dispersed phase = +V V Vd d c( ) [-]
ϕ Flow [m3
/s]
Γ Effective heat capacity [J/K]
µ Viscosity [Ns/m2
]
θ Dimensionless dosing time = t/tdos [-]
ρ Density [kg/m3
]
σ Interfacial tension [N/m]
ζi Yield of component i = ni/nA1 [-]
ζ B0 Initial concentration of nitrosonium ion = 0.035 [-]
Dimensionless groups
Po Power number
Q
N Ddis stirρ 3 5
[-]
Re Reynolds number
ρ
µ
dis stir
dis
ND2
[-]
We Weber number
N Dstir c
2 3
ρ
σ
[-]

Chapter 2
58
Subscripts and superscripts
0 Initial, at t = 0
1 Final (after dosing is completed)
nol Reaction of 2-octanol, see Equation (4)
none Reaction of 2-octanone, see Equation (5)
A Component A (2-octanol)
Aq Aqueous phase (nitric acid solution)
B Component B (nitrosonium ion)
c Continuous (aqueous) phase
C Calibration
cool Cooling
d Dispersed (organic) phase
dis Dispersion
dos Dosing
eff Effective
f Formation
i Component i
max Maximum
Org Organic phase
P Component P (2-octanone)
R Reaction
r Reactor
stir Stirring
w Reactor wall
X Component X (carboxylic acids)
∗ At interface
¯ Average
∞ Ambient

59
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3
Runaway Behavior and Thermally Safe
Operation of Multiple Liquid-Liquid
Reactions in the Semi-Batch Reactor

Chapter 3
64
Abstract
The thermal runaway behavior of an exothermic, heterogeneous, multiple
reaction system has been studied in a cooled semi-batch reactor. The nitric acid
oxidation of 2-octanol has been used to this end. During this reaction 2-octanone
is formed, which can be further oxidized to unwanted carboxylic acids. A
dangerous situation may arise when the transition of the reaction towards acids
takes place accompanied by a temperature runaway.
An experimental set-up was build, containing a 1-liter glass reactor, followed by
a thermal characterization of the equipment. The operation conditions, e.g.
dosing time and coolant temperature, to achieve a high yield under safe
conditions are studied and discussed.
The reaction conditions should rapidly lead to the maximum yield of
intermediate product 2-octanone under safe conditions and stopped at the
optimum reaction time. The appropriate moment in time to stop the reaction can
be determined by model calculations. Also operation conditions are found which
can be regarded as invariably safe. In that case no runaway reaction will occur
for any coolant temperature and the reactor temperature will always be
maintained between well-known limits.
The boundary diagram of Steensma and Westerterp [1990] for single reactions
can be used to determine the dosing time and coolant temperature required for
safe execution of the desired reaction. For suppression of the undesired reaction
it led to too optimistic coolant temperatures.

Runaway Behavior and Thermally Safe Operation
65
3.1 Introduction
To reduce the risk associated with exothermic chemical reactions, in a semi-
batch operation one of the reactants is fed gradually to control the heat
generation by chemical reaction. In practice the added compound is not
immediately consumed and will partly accumulate in the reactor. The amount
accumulated is a direct measure for the hazard potential. A definition of a
critical value of accumulation, to discern between safe and unsafe conditions,
may be rather arbitrary. From a safety point of view an accurate selection of
operation and design parameters is required to obtain the minimum
accumulation.
Hugo and Steinbach [1985] started investigations on the safe operation of semi-
batch reactors for homogeneous reaction systems. Steensma and Westerterp
[1990,1991] studied semi-batch reactors for heterogeneous liquid-liquid
reactions. They demonstrated that it is important to obtain a smooth and stable
temperature profile in the reactor. These authors dealt with single reactions.
However, many problems of runaway reactions encountered in practice are
caused by multiple and more complex reaction systems.
The usual objective is to suppress side reactions whose rates are negligible at
initial conditions but may become significant at higher temperatures, see e.g.
Hugo et al. [1988], Koufopanos et al. [1994], Serra et al. [1997]. In these works
a maximum allowable temperature is defined as the temperature, where
decomposition or secondary reactions are not yet initialized. Limiting the
temperature increase is usually very effective in suppressing side reactions. It is
a rather conservative approach, but necessary to obtain an inherently safe
process, see e.g. Stoessel [1993,1995]. No work has been published on safe
operation of exothermic multiple reactions in which an unwanted reaction is
kept in hand and partially is allowed to take place.
To prevent a runaway one has to operate outside regions of high sensitivity of
the maximum reactor temperature towards the coolant temperature. In case of a
multiple reaction system complications arise: one has to discern between the
heat production rates of the different reactions, see e.g. Eigenberger and Schuler
[1986]. The extension of the theory of temperature sensitivity to multiple, more
complex, kinetic schemes is not obvious: the interaction of parameters in a
multiple reaction system makes the development of an unambiguous criterion
impossible. Each reaction network requires an individual approach and the
optimum temperature strongly depends on the kinetic and thermal parameters of
all the reactions involved.

Chapter 3
66
The present work focuses on the thermal dynamics of a semi-batch reactor, in
which multiple exothermic liquid-liquid reactions are carried out. The runaway
behavior has been experimentally studied for the nitric acid oxidation of 2-
octanol to 2-octanone, and further oxidation products like carboxylic acids. The
kinetics of these reactions have been discussed in Chapter 2. It will further be
evaluated, whether the mathematical model as developed by Steensma and
Westerterp [1990] is sufficiently accurate to predict the reactor behavior and to
stop the reaction at the appropriate moment in time.
3.2 Nitric acid oxidation in a semi-batch reactor
The nitric acid oxidation of 2-octanol to 2-octanone and the further oxidation of
2-octanone to carboxylic acids are described in Chapter 2. The reaction system
was found to be suitable to study the thermal behavior of a semi-batch reactor in
which slow multiple liquid-liquid reactions are carried out. The oxidation
reaction system will be described here briefly.
Figure 1: Schematic representation of mass transfer with chemical reaction
during the oxidation with nitric acid of 2-octanol to 2-octanon and carboxylic
acids.
rnol
rnone
Aqueous nitric acid Phase Organic Phase
2-Octanol2-Octanol
2-Octanone
Carboxylic acids
2-Octanone
Carboxylic acids
Interface

Runaway Behavior and Thermally Safe Operation
67
Reaction system
The oxidation of 2-octanol takes place in a two-phase reaction system: a liquid
organic phase, which initially contains 2-octanol, is in contact with an aqueous
nitric acid phase in which the reactions takes place. The reaction system with
simultaneous mass transfer and chemical reaction is represented with Figure 1.
The oxidation of 2-octanol (A) to 2-octanone (P) and further oxidation products
(X) can be described with the following reaction equations:
A B P B
rnol
+  → + 2 (1)
P B X
rnone
+  → (2)
where B is the nitrosonium ion, which also causes an autocatalytic behavior. The
reaction rates in the acid phase can be expressed on the basis of a second order
reaction:
r k m C Cnol nol A A Org B Aq d= −, , 1 ε1 6 (3)
r k m C Cnone none p P Org B Aq d= −, , 1 ε1 6 (4)
where CA,Org, CP,Org and CB,Aq are the bulk concentrations of 2-octanol (A), 2-
octanone (P) and nitrosonium ion (B) in the organic phase (Org) and Aqueous
phase (Aq), respectively. The kinetic constants knol and knone can be described
with:
k k
E
RT
m HHo= − −%'
()*∞ exp 01 6 (5)
where k∞, E/R and mH0 are the pre-exponential factor, the activation temperature
and the Hammett’s reaction rate coefficient, respectively. H0 is Hammett’s
acidity function, see Rochester [1970]. The value of H0 is plotted as a function
of the nitric acid concentration in Figure 2. The values of the kinetic constants
and the heat effects are listed in Table 1, see also Chapter 2.

Thermally Safe Operation of Nitric Acid Oxidation

Thermally Safe Operation of Nitric Acid Oxidation

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