Calorimetric Investigation of an Exothermic Reaction

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814

Znd Eng Chem. Res. 1994,33,

814-820

Calorimetric Investigation

of

an Exothermic Reaction: Kinetic and

Heat Flow Modeling

R a l p h

N Landau;*+ Donna

G.Blackmond , '

and Hsien-Hsin Tung*

Merck Manufacturing Division and Merck Research Laboratories Merck Company Inc. P.O. Box 2000,

Rahway New Jersey 07065-0900

Util izing an auto mated laboratory reaction calorimeter, an energetic reaction was investigated t o

determ ine feasible pi lot plant operat ing conditions. T he calorimetric data led to kinetic parameters

and a predict ive heat flow model which al lowed simulat ions of the required pilot plant jacket

tem pera ture profiles. On th e basis of thes e profiles, several feasible and safe operat ing condit ions

were determined. Th is study dem onstrates the power of calorimetry as i t relates to process design

an d scale-up of batch an d sem ibatch processes.

In t roduct ion

Pharmaceutical reactions are often accompanied by

significant heat release and mu st therefore be thoroughly

understood t o manage them successfullyon a factory scale.

The heat losses often att end ant to laboratory-scale ap-

paratus make it difficult or impossible to quantify the

rate of he at release associated with the chem istry being

carried out therein. Th e adve nt of computer-co ntrolled,

laboratory-scale calorimeters has provided a tool to

elucidate heat flows accurately to a bout 0.1

W/L

with a

high degree of reproducibility [ l - 4 ] . Combining the

calorimetric data thus obtained with kinetic analysis allows

for the construction of pre dictive heat flow models which

can be app lied to vessels of any volume. T he work

discussed herein focuses on the development of a heat

flow model an d subseq uent simulations of large-sca leplant

vessels ultimately leading t o achievable, safe opera ting

conditions.

Figure

1

shows the desired chemical reaction which

involves th e att ack of a cyclic sulfite by sodium cyanide

in the presence of catalyst and solvent to yield the

hydroxynitrile. This ste p produces an intermediate which

ultimately leads to a drug currently under investigation

in these laboratories.

The proposed plant procedure suggests charging all

components at am bient conditions followed by a ramp at

1.5 C/min from

20

to

90

C where the re action is carried

out to completion. In our laboratory study heat effects

were investigated a t temperatures up t o 110 C.

E x p e r i m e n t a l S e c t i o n

Th e equipment utilized in this study was Mettler's RC1

Reaction Calorimeter 151 outfitted with an MKOl batch

reactor. All com ponents were charged to the reactor a t

ambient conditions followed by

a

reactor temperature

ramp to

90

C at 1.5 C/min. After 15 min a t

90 C,

the

tem peratu re was rampe d at 1.5 C/m in to 100 C. Finally,

after 15 min at 100 C, the temperature was ramped a t 1.5

C/min to 110 C. Th e reactor temperature was held a t

110 C until all therm al activity had ceased. A stirring

ra te of 200 rpm was employed througho ut the experiment.

These conditions were employed t o provide kinetic par-

ameters in th e vicinityof the proposed pilot plant op erating

conditions. Calibrations

to

determine the hea t capacity

* To whom correspondence should be addressed.

t

Merck Manufacturing Division.

1 Merck Research Laboratories.

R

I

Solvent, Catalyst

0,

I NaCN

S

HO

CN

Figure

1.

Desired chemistry which involves the attack of a cyclic

sulfite by sodium cyanide in the presenceof catalyst and solvent to

yield the desired hydroxynitrile.

Reactor

- - - -

Pmbe 0n

I

Time

Figure

2.

Temperature profiles for the jacket and reactor contents

during a typical calibration experiment.

and he at-tran sfer coefficient were performed p rior to an d

after the reaction period.

Ca lo r ime t r i c Procedu re

Th e reaction calorimeter analyzes the raw da ta (tem p-

eratur es and masses) by th e heat flow technique. There-

fore, an energy balance is performe d with th e limits at th e

reactor outer surface. A t any given mom ent in time, for

a batch process, the instantaneous heat flow due to

reaction,

qr,

is given by eq 1, where

U

is the heat-transfer

1)

coefficient, A is the wetted area, Tr is the reactor

temperature,

Tj

is the jacket tem perature, m is the mass

of th e reactor contents, and

C,

is the hea t capacity of th e

reactor contents.

A

calibration experiment imposing

known qr values on the system is used to solve for th e

unknown quantities U and C,. Figure

2

shows the

temperature profiles for th e jacket and reactor contents

during a typical calibration experiment.

Ther e are two distinct parts to th e calibration because

eq

1

has two unknowns. Th e first part of the experiment

involves a short temperature ramp effected by ramping

the jacket temperature. During this period,

qr

is equal to

0 since ther e is no heat flow. Th e second portion of the

calibration involves he ac tion of a precision heating probe

(23.5

W

in this case). Th e calorimeter is instructed to

qr =

UA(Tr

Tj)+

mC,(dTr/dt)

0888-5885/94/2633-0814$04.50/0

1994 American Chemical Society



Ind. Eng. Chem. Res., Vol. 33, No. 4

1994

815

Table 1 Calorimeter Evaluated P arameters

20

110

- -

I

0 5h 100 150

ba

'lime I Minutes

Figure 3. Temperature profiles measured by the automated

calorimeter during the he atu p

and

isothermal period(s).

3

5

l o 1M w

lime

/ h utes

Figure 4. Baseline

and

reaction heat

flows

calculated during the

heat-up and isothermal period(s).

maintain the desired isothermal temperature (using the

jacket to remove the heat flow); thu s dT,/dt is equal to

0

and

q

is equal to 23.5 W. Hence, from these two

experiments, the heat-transfer coefficient

and

the heat

capacity may be determined . Normally, this procedure is

carried out before initiating reactions and afte r reaction

iscompletetoaccoun t forchangesin thesepropertieswith

exten t of reaction. Using these two calibration endpoints,

C,

and UA a t

e

=

0 )

and

C ,

and UA a t

e

=

ef),

the data

analysis allows for several methods of varying (linearly,

proportional

to

heat flow, and others) these properties

over the course of the experime nt. Thi s is necessary since

they generally change during reactions, and eq

1

works

best when the most accurate values for

U

and C, are

employed.

Re s u l t s

Reactor and jacket temperatures were collected con-

tinuously a nd the instantaneous heat flow

was

calculated

over th e course of th e experiment. Figure

3

depicts the

temp erature profiles measured by the calorimeter during

the tem perature ramp and isothermal periods.

Ascan beseen, thereactoreasily maintained thedesired

temperature profile, this being due primarily to i ts

significant heat-transfer coefficient and relatively high

surfaceareatoreactorcontents massratio. Figure4shows

the baseline and reaction he at flows calculated using eq

1

during the tempe rature ram p an d isothermal periods.

Equation

2

relates the total h eat evolved, AHnet, o the

instantaneous heat flow. Th e total heatevolved

was

ound

to be

22.8

kJ (eq

2), or 190

kJ/kg of material charged to

the reactor. In this case,

C,

and

UA were varied in

proportion to the fractional heat evolved. Table 1 ists

heat-transferwff area at t = 0 UA)o 2.17 WIK

heattransfer coeff x area at

t =

if

( U A k 2.72

WIK

1400

J/(kg K)

eat cnpacity (at t

=

0)

C d

heat capacity (at t = if) C d 1750 J/(kg K)

m a

f

MKOl contents 0.120 kg

total

heat released

22.0

kJ

109

'C

diabatic temp rise AT,)

1.25

.. Exp-mt.1 Cddul.*d

0

o

60

l h e fterStanof

Isolhemul

Period/

M u m

Figures. Kineticanalyaisoftheheatflowdataobtainedduringthe

three isothermal reaction periods.

other sign ificant information calculated from th e experi-

ment. Th e final entry in Table

1,

the adiabatic tempera-

ture rise, is the theoretical AT, = AHneJrnCp) emp-

era ture rise which could man ifest itself if cooling capabili-

ties were lost. Th e value reported herein, 109 C,serves

as

an indicator of th e significant exothermic potential of

this reaction. A detailed discussion of these calculations

is available in th e literature [61.

Scrutin y of Figures

3

and 4 reveals th at the three peaks

in h eat flow

(58.5, 78.0,

and 96.6 min) correspond to the

exact times

at

which the reactor temperature reached the

threeiso therm alstage s (90,100,and llO'C, respectively).

Th e subsequen t decay in heat evolution in the isothermal

period following each of thes e peak s is direc tly du e to the

concentration dependence of the rate expression. From

the energy balance, it can be seen th a t the he at flow can

be viewed

as

the p roduct of th e reaction rate, dC /dt, the

specific heat of reaction,

AH-.

and the volume, V, for a

simple reaction, as shown in

eq 3.

Th e implication of eq

(3)

3

for this experiment is that three isothermal kinetic

experiments are embedded in these data. Th e decay in

hea t flow during each isotherm al period is proportional

to

the expected decay in reaction rate, dC/dt, during th at

time. A pseudo-first-order model fo r reaction is given in

eq 4, where Co and qa correspond to the concentration

dC/dt = -kC- ln(C/Co) = -kt - n(qJq,J = -kt (4)

and heat flow values at the beginning of an isothermal

period. Th e relationship depicted in eq 4 allows for

conversion of the heat flow data obtained with the

calorimeter into kinetic information aspseudo-first-order

rate constants. Figure 5 shows the results of such an

analysis whichappear to fit a

pseudefirst-order rate

model

well.

Table

2

lists the significant statistical parameters

obtained during th e regression analysis.

For predictive purposes, it would be useful

to be

able

to model the rate constants' variation with temperature.

Equation 5 is the well-known A rrheniusexpression, w here

(5)



816 Ind. Eng. Chem. Rea., Vol.

33,

No.4,1994

Table 2. Kinetic Regression Analysis

R e su l t s

T 'C) k (min-1) stand . error (m in-9 correln coeff

9)

90

0.0016

O.ooo4 0.951

100

0.0328

O.ooo4 0.998

110

0.0737

O.ooO8

0.987

-25

-3

0ExperinenW

-Arrhdus

-5

a m a m 5 a m a m 0.Mu)

1fl

OK)

Figure

6. Arrheniw plot of the ratc const nts obtained from th e

three isothermal reaction periods

Table 3. Arrhenius

Remnsi ion

Analysis

R e su l t s

param value stand . error

E

31 kcaVmol

5

kcaVmol

Ao 6.0 x 10 min-L

2 x 10

min-1

k is the rate con stant, A0 is the preexponential factor,

E.

is the activation energy, R is the gas law con stant, and

T

is the temperature. Figure 6 depicts the fit of th e rate

constants given in Table 2 to the Arrhenius expression (In

k vs lln. The Arrhenius expression fits the data

reasonably well 9= 0.978), lthough only three points

were available for the regression. Ta ble 3 lists the

significant statistical param eters.

Th e use of only three isothe rmal estimates of th e rate

constant to determine A rrhenius parameters would seem

insufficient. However, it is the p urpo se of this stu dy to

predict the global heat evolution rate (not the individual

species concen tration profiles) in order

to

carry out the

reaction safely on a larger scale. Therefore, slight devia-

tions from linearity (as een in Figure 6) are acceptable.

For the current o bjective, a polynomial could be used to

describe the rate constants' variation with temperature;

however, th e more fam iliar Arrhen ius expression was used.

Th e hea t flow modeling section will show tha t th e slight

deviations from linea rity have little or no impact on the

predictions. It should be clear from the procedu re

discussed previously that an elegant, however simple,

app roac h for the de term inatio n of pseud o-first-order (or

other orders via insertion into eq

4)

rate constants is

possible.

Proceeding with this analysisrequires tha tth e chemical

pathways do not vary significantly over the temperature

range. Th is assum ption is valid providing th at ultimate

tem pera tures do not exceed the maxim um observation of

110 C.

Equation

6

elates the fractional heat evolution, FH E,

to the instantaneo us he at flow, qr. Th e only difference

between the two integrands is the upp er limit of integra-

tion, and by th is definition, FHE always varies between

0

and 1. Due

to

the re lationships elucidated in eqs 3,4,

and 6 he fractional he at evolution, FH E, may be loosely

(in

the case of incomplete reactions)

or

precisely (in the

case of comp lete single reac tions) related to the extent of

reaction,

or

conversion. Figure

7

shows theFHEcalculated

Figure 7. Fractionalheat evolution

ealeulatad

with eq 6 during the

heatu pand isothermalperiod(s).The AFHEpericdsshoamindieate

the fractional amoun tof heat r e l e d during the isothermal periods

at

90

nd 100

'C.

lkrl

onrlnion

H 1 G e m d o n k l n y h n i

1 . f l l W

$ 1

:

5

V O

E 0 5

0 1 0 0 m 3 0 0 1 1 0 0 5 0 0 6 c m

T i e M i n u t e

Figure

8.

Predicted heat

f low

expected for propod operating

cond itions. Temp erature ramped from 20to

90

C at 1.5 C/rnin,

then held at

90

C .

with eq 6 during the temperature ramp and isothermal

periods. Figure 7 also depicts the regions in time when

the system was holding at th e

90

nd 100 'C isothe rms ,

and th e point a t which the system reached 110 C.

I t is interesting to note th at all of these he at flow and

kinetics parameters describing a complex and energetic

reaction were obtained from two simple temperature

measurem ents of th e reactor and the jacket in a non

invasive in sit u experiment. These parameters can now

be used to develop a predictive model for other reactor

vessels and operating con ditions.

Heat Flow

Modeling

Using the regressed Arrhenius parameters for any

assumed temperature profile for the reactor contents, a

reasonab le estimate of th e rate of hea t evolution may be

made. Clearly, a t any given tempe rature, a rate constant

maybe calculatedfromeq5 and theArrheniusparameters

previously obtained. From this, a prediction of the

expected he at flow can be made.

As

a starting point, the

predicted heat flow diagram for the proposed operating

conditions is shown in Figure 8. For the conditions in

Figure 8, the peak in the heat generation curve occurs

when thebatchfirstreaches9 C. Additionally,themodel

predicts a maximum heat flow of 3.1

W

for the laboratory-

scale reactor upon initially reaching 90 C.

Equation

7

relates th e predicted m ass specifichea t flow,

&(heatflowdivided

bymass),toanadiabatictemperature

rise rat e (dAT.d/dt). Using th e initial heat capacity given

(7)

in Table

1,

1400

/ (kg

K),

he maximum predicted heat



Ind . Eng. Chem. Res., Vol. 33, No.

4, 1994

817

0

40 O

120

lko

200

Tune Minutes

Figure 9.

Predicted heat flow expected for modified operating

conditions. Temperature ramped from 20 to 110 OC at 1.5 OC/min,

then held at 110 OC.

flow, approxim ately

1.5

kJ/(kg.min), converts to

1.07

C/

min adiabatic temperature rise rate with eq

7.

Pilot plant

and factory-scale vessels typically displace approx imately

500-3000

gal for batch an d semibatch processes. These

vessels can d issipate he at a t approximately

0.02

C/min

when operating without the aid of forced cooling [A.

Therefore, the temperature rise rate,

1.07

C/min, is

significant, and could lead t o a runaway reaction.

Th e mass specific hea t flow can also be con verted back

into an ordina ry hea t flow simply by multiplying by t he

total mass being employed. For this experiment, th e total

mass was

0.120

kg, which converts he maximum predicted

he at flow of 1.5 kJ /(kgmin) to 0.18 kJ/min or 3.1 W (as

shown in Figure 8). Referring to Figure

4,

this com pares

well to the value of 3.7 W obtained experimentally.

Figure

9

shows th e predicted hea t flows for a modified

reactor tempe rature profile where th e initial ram p a t 1.5

C/min to 90 C was extended to 110 C.

Again, using th e initial heat capacity given in Tab le

1,

1400 J/(kg K), the maximum predicted heat flow, ap-

proximately

9

kJ/(kgmin), converts to

6.43

C/min

adiabatic temperature rise rate w ith eq 7. Clearly, this

adiabatic tem pera ture rise rate is significant, and it will

be shown in the ne xt section th at this will lead to a runaway

reaction under pilot plant or factory conditions. For th e

experiment, the maximum predicted heat flow of

9

kJ/

(kgm in) becomes 18 W. Referring to Figure 4, this exceeds

the value of

9

W obtained in the stepped temperature

profile obtained in th e experimental data, due for the most

par t to th e consumption of reactan ts during the waiting

periods at 90 and 100 C in our experiment. Thi s claim

will be substa ntiated in the following. If the fractional

am oun t of he at released is considered (Figure

71,

th e model

and th e experimental data would be nearly identical. From

the corresponding changes in the FH E, an estim ate of th e

FH E which would have been observed (ha d he calorimeter

gone directly at 1.5 C/min to 110 C) is possible. Th e

calorimeter reached 110 C at approximately 96.6 min,

when the FH E was

0.70.

Subtracting the changes in FH E

durin g the two prior isotherm al periods gives

0.70 0.22

-

0.14

=

0.34.

Th e model also calculates conversion (as

seen in Figures 8 and 9) and can be directly compared to

the estimate provided by th e calorimeter data. Figure

10

shows the predicted conversion along with the cor-

responding tem pera ture profile. Th is simulation star ted

a t 20 C, ramped at 1.5 C/min up

to

110 C, and then

held at

110 C.

At the mom ent the reactor temperature

reached

110

C, th e conversion was

0.37.

This compares

well to th e estimated value of

0.34

which would have been

obtained by the system had it gone directly to

110

C.

Th e initial conclusion is that the model does a reasonable

job of estimating th e rate

of

reaction and the rate of heat

evolution. Th e model could be improved by exploring

30 35 40 45

50 55

60

65

Time Minutes

Figure10.

Predicted conversionand temp erature duringsimulation

of heat flow for a reactor temperature profiie rampingto 110 O C at

1.5 OC/min and holding. Th e model indicates that the Conversion

is

0.37

at the time 110 C

is

reached.

more complex reaction rate expressions. Th e objectives

of thi s work simply requ ire a model providing a reasonable

estimate of reaction rate, and therefore heat evolution

rate, which may be used

to

provide additional insight

to

help identify an intrinsically safe operating procedure.

Pilot Plant Modeling

Th e laboratory d ata o btained has been utilized toarrive

a t a kinetic model of the specific hea t flow. Th e ultimate

use of th e model thu s developed is

to

predict plan t vessel

performance. Th e ability

to

predict such behavior depen ds

primarily on our es tima tes of th e intrinsic properties of

th e pla nt vessel, namely, th e overall heat-transfer coef-

ficient. Th e influences of changes in reactor geometry,

agitation, and mass transfer when moving from the

laboratory

to

the pilot plant have been assumed

to

be

minor relative

to heat-transfer issues. Th e laboratory

reactor provides bette r mixing, and th erefore its observed

reaction rate will be equal to or higher than th at expected

in the pilot plant. Th e following simulations then

repre sent a worst-case scenario.

Observ ations of th e proposed pilot plan t vessel indicate

th at the best consistent cooling rate is approximately 1.5

C/min while maintaining a

75

C

A T

across th e jacket

durin g a dummy ru n (solvent charge only, therefore qr =

0).

Using eq 1, an estimate of th e effective

UApaot plant

can be obtained. Solving eq 1 for

UA,

and using

C, =

1400

J/(kg K ) (as obtained in the calibration) and m =

118.7

kg (the mass of solven t used in th e

runs , UApdot

plant is

found to be

3324

(J/min)/K or

55

W/K. The normal pilot

pla nt charge is 154.7 kg; therefore UApilotplant is expec ted

to increase since the w etted a rea will also increase, and in

the absence

of

other information, the mass ratio

(154.7

kg/118.7

kg- 1.303)willbeusedtoobtainanewUAmotpbt

of

71.7

W/K. This

UApaot plant

may be used to simulate

several scenarios in th e pilot plant, including th e desired

operating conditions, altern ate tem peratu re programming

options, an d investigation of sem ibatch approaches.

Th e relative heat-transfer ability of th e laboratory and

the pilot plan t vessels may now be com pared. Tab le

1

listed the UA prior to

and

after the reaction period for the

laboratory reactor, with an average of

2.45

W/K. To

compare this to the pilot plant value of

71.7

W/K , both

mu st be divided by th eir respective mass charges, yielding

(

U A / ~ ) M K ~ ~20.4(W/K)/kgand (UA/m)pd0t

p h t

=

0.463

(W/K)/kg. These figures indicate that the laboratory

vessel transfers heat a t 44 times the ra te of th e pilot plant

vessel per un it mass present. Therefore, it is clear tha t

while laboratory studies may indicate ease in handling

heat removal, often larger vessels

will

present significant

challenges in operating a t similar conditions.

Using the model developed earlier, estimates of the

required jacket tem peratu re may be made using eq 8, which



818

Ind. Eng. Chem. Res., Vol. 33, No. 4, 1994

-50-

-100

is a rearrang emen t of eq 1. Equation8 eveals that, during

temperature ramping, the dTr/dt term will reduce the

required cooling and t ha t the maximum cooling will occur

when the reactor tem peratu re initially reaches ita desired

set point, since from th at p oint forward, the dT r/dt term

is zero.

Figure 11shows th at t he required jacket temp erature

profile for the curre nt proposed operating procedure (begin

a t 20 C, ramp at 1.5 C/min to 90 C, then hold at 90 C)

calls for a rapid decrease in temp eratur e w ithin a narrow

window of time . Figure 1 2 shows the same simulation,

for

conditions where the temperature ram p for the reactor

continued up to 100 C and then held at 100 C. It is clear

th at once the reactor reaches 100 C, there is no hope of

regaining control since no conventional cooling appara tus

may be expected to carry out the almost 200 C decrease

in the short time required. Th e calorimetry, heat flow

modeling, and plant modeling clearly indicate tha t th e

result of such an overshoot w ill be a d angerou s runaway

reaction.

Analysis of the required jacket temperature response

curves presented in Figures 11and 1 2 ndicates th at several

factors should be considered w hen seeking an intrinsically

safer method to run the process while maintaining its

productivity. Practically, th e discontinuity in the slope

of the

T,

curve, dTr/dt, at the moment the reactor

temperature reaches t he desired set point

90

or 100 C,

depending on the scenario) causes extreme difficulty as

displayed in th e acket temperature response curves. Th e

solution is to con sider a m ore realistic reactor temp eratur e

profile. Initially, dT r/dt will be equa l to 0; hen, as the

jacke t tem pera ture rises, dT,ld t will slowly increase. As

the tem perature ap proaches the upper set point, it would

be desirable for dTr/dt to tend toward

0

once again.

Mathematically, this is shown in eq

9.

Reactor acket

1

Equation 9 suggests the use of a parabo la to repres ent

the derivatives' variation with time since it mu st pass

through

0

a t both ends of the ramping period, and i t has

a maximum. Using an expression suc h as eq 10 for

2 5 -1 /

0

d

Tr

a t 2

+

bt dT,

=

J ( a t 2

+

bt) dt ==+

d t

Reactor acket

T, =

a t 3 + g +

c (10)

3 2

obtaining a reactor tem pera ture profile will significantly

reduce the stiffness of th e required jacket temp erature

response. Because th e derivative is

0

a t t

=

0,

there is no

zero-order parame ter in the expression for dTr/dt. Th e

initial tem pera ture is assumed to be 20 C; therefore the

con stant of integration

c

in thi s case) will be e qual t o 20.

Figure

13

shows this app roach qualitatively. Th e value

of dTr/dt reaches a maximum, exactly half way to the

desired upper se t point for the reactor temperature during

the tempe rature ramp. Using this approach, there are no

sudden changes in d Tr/dt a s there were in th e previous

examples. Th is will provide th e safest mode for a fully

batch operation. Th e variable which determines th e

constants

a

and b in eq 10 is the t ime taken to reach 90

C. Realistically, under p ilot plant conditions, several

linear ram ps would likely be executed to simulate the

parabolic dT,/dt constraint.

Reactor acket

0 -

0 250 500 750

11

Jo

Time

/

Minutes

Figure

11.

Required acket temperatureprofile

to

maintain desired

reactor temp erature profile (ramped from 20 to

90

C at

1.5

OC/min,

then held at 90 C).

Time

Figure

13.

Representation

of

a parabolic dTJdt curve which resuita

in a smooth approach to the desired upper set point for the reactor

temperature.

0

Time

I Minutes

Figure

14.

Required jacket tem perature response for a parabolic

dT,/dt constraint employing a 120-min heat-up time.

Figure 14 shows the resulting required jacket temper-

ature profile if the time taken to reach 90 C is 120 min.

Th e relationship for the reactor temperature profile is





820 Ind. Eng. Chem. Res., Vol. 33,

No.

4 1994

'lime

Minutes

Figure 18. Predicted heat flow shown for

semibatch operation

with

a

180-min addition

period and reaction

at

90OC

if a situation should arise which suggests th at t he reaction

is beginning to run away, an adverse situation may be

averted by simply stopping th e feed. Th e heat flow will

decay, control will be regained, and th e add ition can be

resumed. In essence, runn ing in

a

semibatch mode will

provide two leverages against a runaway reaction: th e

jacket temperature and the feed rate.

Conclusions

Th e work presented herein de monstrates the de pth of

useful an d predictive information which may be obtained

from a simple set of calorimetricmeasurements. A kinetic

and heat flow model was developed from these m easure-

me nts which was applied to th e scale-up of heat-tra nsfer

problems in batch and semibatch processes leading to

intrinsically safe, feasible operating conditions. In pa r-

ticular, the power of automated laboratory calorimeters

was proven

as

simple empe rature data obtained from this

instrument were the basis for all of t he d eterminations

made.

Referring t o the example cited , two safe and feasible

alternatives for pilot plant operation were discovered one

which allowed for a fully batch mode of operation and

ano ther which considered a semibatch alternative. This

example illustrates th e need to integrate developmentand

safety objectives into a common goal, ultimately yielding

highly productive, reliable p rocesses.

As a final note, it should be understood th at caution

should be exercised in the ap plication of these concepts.

For example, in the case stud y presente d herein, initially,

a parabolic dT,ldt constraint up to

70

C or the first

batch would be approp riate to

test

th e model safely. Once

confidence is gained with the model and pilot plant

performa nce, a more aggressive approach m ay be take n.

Nomenclature

= time, min

qr =

instantaneous heat

flow

W

qa

= instantaneous heat

flow

at t = 0, W

U = overall he at-transfer coefficient, W /(m2K)

UO overall heat-tran sfer coefficient at = 0, W/(m2 K)

Ut= overall heat-transfe r coefficient at t = tf , W/(m2K)

A = area available for heat transfer, m2

A

=

area available for heat transfer a t

=

0, m2

At = area available for heat transfer a t

t =

tf, m2

TI temperature of reactor contents, C

dTI/dt = time derivative

of

reactor contents temp erature,

Tj = temperature of jacket contents, C

m = mass of reacto r contents, kg

C, =

hea t capacity of reactor con tents , J/( kg

K)

t =

extent of reaction

AHnet

= total heat released or consumed, kJ

AT,d

=

adiabatic temperature rise, C

C

=

conc entration, mol/L

CO

concentration at t

=

0, mol/L

dC/dt

=

general rate of reaction, mol/min

A , = molar heat of reaction, kJ/mol

V

= volume of rea ction mass, L

k

= pseudo-first-order ate constant, min-1

A = pree xpo nen tial factor in Arrhenius expression, min-l

E,

= activation energy in Arrhenius expression, kcal/mol

FHE =

fractional heat evolution

Qr

=

mass specific heat flow, W/kg

UApilot

p k t = overall hea t-transfer coefficient X wetted area

a =

third -order coefficient in temp erature expression, C/s3

b

=

second-order coefficient in temp erature expression, C/

c

=

zero-order coefficient in temperature expression,

C

m,= mass of rea ctor co nten ts, kg

C,, = heat capacity of reactor co ntents, J/(k g K)

dmddt

=

mass addition rate of dosed material, kg/s

C,

= heat capacity of adde d mate rial, J/( kg K)

T d =

temperature of added material, C

U s

for pilot plant, W/K

92

Literature Cited

(1) Regenass,

W. Thermal and Kinetic DesignData from

a Bench-

Scale Heat

flow Calorimeter.

ACS Symp. Ser. 1978,65, 37.

(2)Karlsen, L. G.; Soeberg, H.; Villadsen, J. Optimal Data

Acquisition

for Heat Flow Calorimetry. Thermochirn. Acta

1984,

72, 83.

3) Litz,

W.

The Thermokinetic Reactor TKR and ita

Possible

Applications n Chemical Research and Engineering. J . Therm. Anal.

1983, 27,

215.

(4)

Landau,

R.

N ;

Williams,

L.

R. Reaction

Calorimetry:

A

Powerful

Tool

Chem. Eng.

Prog. 1991,87 (12), 65-69.

(5)

Mettler Instrument Corporation,69

Princeton-Heightatown

Rd., P.O.

Box

71,Heightatown, NJ 08520-0071.

6) Landau R.

N.;

Cutro, R. S. Assess Risk Due to Desired

Chemistry. Chem. Eng. Prog.

1993,89

(4), 66-71.

(7)Sharkey, John J.; Cutro, Robert S.; Fraser, William J.;

Wildman, G eorgeT rocess Safety Testing Program for

Reducing

Risks Associated With

Large Scale Chemical Manufacturing Opera-

tions. Presented a t

the AIChE 1992Loss Prevention Symposium,

March

31, 1992.

(8)

Mills, P.L.;

Ramachandran,

P.

A.; Chaudari,

R.

V.

Multiphaee

Reaction Engineering

or

Fine

Chemicals

and

Pharmaceuticals.

Rev.

Chem.

Eng.

1992,8 (1-2), -176.

(9)Paul,

E. . Design

of

Reaction Systemsfor

Specialty

Organic

Chemicals. Chem. Eng.

Sci. 1988,43, 1773-1782.

Received

for

review September

2,

1993

Revised

manuscript

received

December 27, 1993

Accepted January 14, 1994.

* Abstract published in Advance

ACS

Abstracts, March 15,

1994.

Documents

Calorimetric Investigation of an Exothermic Reaction