The oxidation of sewage sludge in the liquid water phase atelevated temperatures and pressures : wet-air oxidationCitation for published version (APA):Ploos V Amstel, J. J. A. (1971). The oxidation of sewage sludge in the liquid water phase at elevatedtemperatures and pressures : wet-air oxidation. Eindhoven: Technische Hogeschool Eindhoven.https://doi.org/10.6100/IR114081

DOI:10.6100/IR114081

Document status and date:Published: 01/01/1971

Document Version:Publisher’s PDF, also known as Version of Record (includes final page, issue and volume numbers)

Please check the document version of this publication:

• A submitted manuscript is the version of the article upon submission and before peer-review. There can beimportant differences between the submitted version and the official published version of record. Peopleinterested in the research are advised to contact the author for the final version of the publication, or visit theDOI to the publisher's website.• The final author version and the galley proof are versions of the publication after peer review.• The final published version features the final layout of the paper including the volume, issue and pagenumbers.Link to publication

General rightsCopyright and moral rights for the publications made accessible in the public portal are retained by the authors and/or other copyright ownersand it is a condition of accessing publications that users recognise and abide by the legal requirements associated with these rights.

• Users may download and print one copy of any publication from the public portal for the purpose of private study or research. • You may not further distribute the material or use it for any profit-making activity or commercial gain • You may freely distribute the URL identifying the publication in the public portal.

If the publication is distributed under the terms of Article 25fa of the Dutch Copyright Act, indicated by the “Taverne” license above, pleasefollow below link for the End User Agreement:www.tue.nl/taverne

Take down policyIf you believe that this document breaches copyright please contact us at:openaccess@tue.nlproviding details and we will investigate your claim.

Download date: 17. May. 2020

THE OXIDATION OF SEWAGE SLUDGE IN THE LIQUID WATER PHASE

AT ELEVATED TEMPERATURES AND PRESSURES

(WET-AIR OXIDATION)

Fl t,;l-IT WATER

'POLLUTION NOW

J.J.A. PLOOS VAN AMSTEL

THE OXIDATION OF SEWAGE SLUDGE IN THE LIQUID WATER PHASE

AT ELEVATED TEMPERATURES AND PRESSURES

(WET-AIR OXIDATION)

PROEFSCHRIFT

TER VERKRIJGING VAN DE GRAAD VAN DOCTOR IN DE

.TECHNISCHE WETENSCHAPPEN AAN DE TECHNISCHE HO­

GESCHOOL TE EINDHOVEN, OP GEZAG VAN DE RECTOR

MAGNIFICUS, PROF.DR.IR.A.A.TH.M.VAN TRIER,VOOR

EEN COMMISSIE UIT DE SENAAT IN HET OPENBAAR TE

VERDEDIGEN OP VRIJDAG 2 APRIL 1971 TE 16 UUR

DOOR

JOHANNES JACOBUS ASUERUS

PLOOS VAN AMSTEL

GEBOREN TE EINDHOVEN

DIT PROEFSCHRIFT IS GOEDGEKEURD DOOR DE PROMOTOR

PROF. DR. K. RIETEMA

CO-REFERENT

PROF. DRS. H.S. VAN DER BAAN

AAN OE KOMMUNE NUENEN

EN IN HET BIJZONOER AAN

PRINSES LISELORELEI

ACKNOWLEDGMENT

The thesis in this make-up would not have been completed but

for the contribution of many whom I would like to thank here.

I will mention in particular Mr. J. Bos for his interest in

the project and for his advice and assistance in the research

of the literature.

Thanks are also due to Mr. W.C. Koolmees who advised on the

design and construction of the apparatus, and to Messrs. P.A.

Hoskens and A.H. van der Stappen, who constructed and put to­

gether the equipment.

Furthermore I am indebted to Messrs A.W.C.M.van Alphen, F.C.M.

van de Berg, C.G.M. de Boer, F.H.J. Bukkems, P.J.A.M. Derks,

H. van Gool, G. Groen, c. van de Moesdijk, P.J.T. Samuels ' H.J.C. Slegers, F.C.R.M. Smits, H.P.E. van de Venne and P. van

Zutphen, who carried out most of the experiments.

Almost all technical drawings have been made by Mr.J.Boonstra,

the remainder by Mr.Klein Wassink. Mrs. D.M.Vermeltfoort typed

the text which was edited by Miss G.M. Kurten. I would like to

thank both ladies and both gentlemen for the accuracy with

which they carried out their work.

I am very grateful to Mr. H.J.A. van Beckum, who made this

thesis readable by correcting the language and to Ton Smits,

the famous artist,who made this thesis digestible by his witty

cartoons.

I wish to thank in particular the directors of the Architekten

en Ingenieurs Bureau of the N.V. Philips' Gloeilampenfabrieken

for the assistence offered during the final stages of this

work.

Finally, I include in my acknowledgement my wife and also

Mr. J. Waterman for their encouragement during the last weeks

of the preparation of the thesis.

VI

CONTENTS

SUMMARY

1. TREATMENT AND DISPOSAL OF SEWAGE

2.

1.1. INTRODUCTION

1. 2. WASTE WATER TREATMENT

1.3. SEWAGE SLUDGES

1.4. SEWAGE SLUDGE TREATMENT AND DISPOSAL

THE WET-AIR OXIDATION PROCESS

2.1. EVOLUTION OF THE PROCESS

2.2. OXIDATION "ROUTES"

2.3. INFLUENCE OF PROCESS PARAMETERS

2.4. END PRODUCTS

2.5. COST OF PROCESS

2.6. LITERATURE REVIEUWS

2.7. CONCLUSIONS

x

1

1

2

3

5

8

8

9

10

13

13

14

14

3. THE OXIDATION OF A GLUCOSE SOLUTION AS A MODEL SLUDGE 15

3.1. PRELIMINARY EXPERIMENTS

3.1.1. Apparatus and experimental details

3.1.2. Thermal treatment and its influence on

the oxidation

3.1.3. Influence of shaker frequency

3.1.4. Homogeneous oxidation of glucose

3.1.5. The maximum conversion

15

15

16

19

21

23

VII

3.1.6. Mathematical description of conversion

rate

3.2. CONTINUOUS FLOW EXPERIMENTS

3.2.1. Apparatus and experimental details

3.2.2. Influence of temperature

3.2.3. Influence of COD concentration in the

24

25

25

27

feed and oxygen pressure 29

3.3. THEORETICAL ANALYSIS OF RESULTS AND DISCUSSION 30

3.3.1. Introduction

3.3.2. Model for the macro kinetics

3.3.3. Model for the micro kinetics

3.3.4. Speculations on the order of magnitude

30

32

36

of kinetic data 40

4. THE DISSOLUTION OF SLUDGE

4.1. APPARATUS AND EXPERIMENTAL DETAILS

4.2. THE HYDROLYSIS OF SLUDGE PARTICLES

4.2.1. The rate of hydrolysis of activated

sludge

46

46

47

47

4.2.2. The rate of hydrolysis of primary sludge 51

4.2.3. Repeated-hydrolysis 53

4.3. INFLUENCE OF CONCENTRATION AND OF OPERATING

CONDITIONS ON [cmax] 54

5. THE OXIDATION OF PRIMARY SLUDGE 57

5.1. PRELIMINARY BATCH EXPERIMENTS

5.2. SEMI-BATCH EXPERIMENTS

57

59

VIII

5.2.1. Apparatus and Experimental details 60

5.2.2. Model for oxidation in a semi~batch system 62

5.2.3. Results and discussion 66

6. THE OXIDATION OF ACTIVATED SLUDGE

6 .1.

6.2.

6.3.

6.4.

EXTENSION OF THE MODEL

APPARATUS AND EXPERIMENTAL DETAILS

EXPERIMENTS WITH SLUDGES OF DIFFERENT ORIGINS FURTHER EXPERIMENTS WITH EINDHOVEN SLUDGE

6.4.1. Influence of temperature

74

74

76

77

79

79

6.4.2. Evaluation of kinetic data and discussion 81

6.4.3. Influence of pressure 86

7. GENERAL DISCUSSION AND CONCLUSIONS 89

7 .1.

7.2.

7.3.

COMPARISON BETWEEN MODEL SLUDGE AND OTHER SLUDGES

7.1.1.

7.1.2.

Reaction orders

Locus of oxidation and diffusion

limitation

7.1.3. Reaction rate constants

7.1.4. Effect of dissolution

7.1.5. Sludges of different origins

ANOMALOUS PHENOMENA

CALCULATIONS OF THE SIZE OF COMMERCIAL REACTORS

REFERENCES

NOMENCLATURE

SAMENVATTING

LEVENSBESCHRIJVING

89

89

90

90

92

93

93

94

102

107

110

113

IX

SUMMARY

Wet-air oxidation is a method of sewage sludge treatment by

which the sludge is oxidised in a liquid water phase in the 0 presence of air at temperatures of about 200-300 c and pres-

sures of 40-120 atm. From an analysis of the literature on the

subject it became clear that this process had apparently been

developed empirically and that only little insight into the

fundamental aspects was present.

Because of the attractiveness of wet-air oxidation, a research

project was carried out in which a process engineering ap­

proach was applied.

First the oxidation of a glucose solution as a model sludge

was investigated with semi-batch and continuous flow experi­

ments at about 200°c and 50 atm.It followed from these experi­

ments that the oxidation of the model sludge was fast compared

with diffusion of oxygen. This caused the oxidation to take

place within the diffusion layer around the gas bubbles. The

chemical reaction rate can be described as first order in or­

ganic matter and zero order in oxygen, while the reaction rate -1 0

constant is about 2 sec at 200 c.

A model is presented for the conversion in the continuous flow

reactor, including combined reaction and diffusion, mixing and

convection. This model gives a fair description of the influ­

ence of process parameters on the conversion and of the con­

centration profiles in the reactor.

At the high temperatures applied in practice real sludges

partly pass into solution by hydrolysis. Owing to this, simul­

taneous oxidation of hydrolised sludge and sludge particles

takes place.

In order to understand the contribution of the oxidation of

hydrolised sludge to the total conversion rate, the hydrolysis

of sludge was investigated with batch experiments.

x

The effect of hydrolysis on the overall conversion rate was

studied, using primary sludge. Oxidation experiments were

carried out at 230°c and 100 atrn with sludge as such, with

hydrolised sludge and with a suspension of the solid residue

of hydrolysis.

From these investigations it was concluded that hydrolysis

does not influence the overall conversion rate, since hydro­

lised sludge and solid sludge have almost the same reactivity.

The oxidation of activated and primary sludge proceeds more

slowly than the oxidation of the model sludge, and the conver­

sion takes mainly place in the bulk of the sludge. The degree

of diffusion limitation of oxygen, or the extent to which the

conversion rate is reduced by oxygen transfer depends on temp­

erature. At 180°c mass transfer can be neglected, while at

290°c the conversion rate is largely determined by the rate of

mass transfer.

A model for the oxidation of sludges is presented.The starting

point of this model is that in the sludge two groups of com­

ponents can be distinguished which differ in reactivity and

which are oxidised simultaneously, while furthermore a third

group is present which is completely inactive. Experiments

have shown that activated sludge consists for about 65% of

more reactive matter, 25% of less reactive matter and 10% of

inactive matter. The chemical reaction rate was described as

first order in organic matter and first order in oxygen (at

relatively high oxygen pressures, the reaction becomes zero

order in oxygen). Mass transfer is also included in the model.

It was found that this model provides a fair description of

the experimental findings. The effect of temperature on the

reaction rate constants of both oxidisable groups in activated

sludge can be described by Arrhenius' law,while the activation

energy is about 23 kcal/mol for both groups. At 255°c the re­

action rate constants are 250 and 20 m3/kg h, respectively.

XI

On the ground of the results of the research the size of a

commercial reactor was calculated. Depending on the amount of

surplus oxygen, the required residence time in a plug flow re­

actor operated in co-current was two to three times as long as

in a plug flow reactor operated in counter-current, while the

required residence time in a reactor in which the liquid was

completely mixed, was 6 to 12 times as long as in a counter­

current reactor.

In practice, a co-current reactor is applied, in which the

mixing state will be somewhere between completely mixed and

plug flow, which will require a residence time of about five

times that in the counter-current plug flow reactor. For prac­

tical reasons, however, the co-current reactor applied in com­

mercial installations, will probably remain more attractive.

XII

-

~,,, '

Chapter 1 TREATMENT AND DISPOSAL OF SEWAGE

1.1. INTRODUCTION

For many centuries sewage has been discharged into streams, lakes and ponds, and even now this is a very normal procedure. In these natural waters micro-organisms consume the discharged

contaminations, while at the same time the solids of the sew­

age settle, both mechanisms resulting in natural purificat.ion. In order to avoid pollution of natural waters, the discharge

of sewage should balance the natural purification capacity of

the receiving water. When the flow of waste water exceeds this

capacity, the contaminations in the waste water must be re­duced by a suitable treatment.

In general, the pollution of natural waters is objectionable

for the possible hazard to public health and safety.Of a less­er consequence, but still very real, is the aesthetic aspect of the deterioration of surface water. Before long this aspect

1

will also weigh more heavily, because the increasing amount

of leisure time and wealth imply increased recreation at, on

and in the water (1, 2, 3).

For the reduction of contaminations in the waste water a varie­

ty of waste water treatment processes is available ( 1, 2, 4,

5, 6, 7). Such a treatment generally results in a flow of more

or less clean water and in a second small flow containing con­

centrated suspended impurities known as sewage sludge. In the

present procedures the sewage sludge undergoes a further treat­

ment towards a form suitable for one or another method of fi­

nal disposal, like dumping or soil conditioning (11, 33, 34,

35, 36).

1.2. WASTE WATER TREATMENT

Figure 1.1 shows a conventional activated sludge installation.

sewage

Figure 1.1.

A B

grit

D

to surf ace

water

air activated 1----~'--~~~~~~~--' sludge

primary sludge

Sewage treatment plant.

A: Sereening; B: grit removal; C: primary settling;

D: aeration; E: seeundary settling.

2

The treatment of domestic sewage and many other organic waste

waters may usually be divided into two steps: pretreatment and

biological oxidation. Pretreatment includes screening, grit removal and sedimentation or flotation. The sludge removed

from the settling or flotation tanks is called primary sludge.

The pretreated waste water can be further subjected to biolog­ical oxidation, resulting in a removal of colloidal and dis­

solved organic matter by the action of micro-organisms. In the

activated sludge system the pretreated waste is brought into

contact with the activated sludge, which consists of floccu­

lated micro-organisms and adsorbed contaminations, the process being carried out in an aerated tank.

The mixture of activated sludge and treated water is subjected

to a .secondary sedimentation. The activated sludge is partly recycled to the aeration tank, maintaining stationary condi- 1

tions. The surplus sludge is removed for further treatment andi disposal.

The overflowing liquid of the secondary sedimentation tank is . discharged into the natural waters or used to irrigate the .

land. At present another step is sometimes added to reduce the nitro­

gen and phosphorus content (8, 9, 10).

1.3. SEWAGE SLUDGES

Municipal sewage consists of aqueous discharges from kitchens,

bathrooms, lavatories and laundries, and also of waste waters

from a variety of industries. Since primary sludge consists of the settable contaminations of the sewage, its composition depends on the habits of the population and on the kind of industry discharging on the mu­

nicipal sewerage. A typical composition of the organic matter in a suspension of

primary sludge is as follows.

Protein Lipids Starch Crude fibre Volatile acid Total

kg/m3 6.2 10.5 3.6 13.5 1.3 35.1

3

Deviations from these figures have to be expected. The amount

of inorganic matter in primary sludge is also subject to vari­

ations and may differ from plant to plant. The primary sludge

of Eindhoven contains an amount of inorganic matter of about

20 kg/m3.

The composition and physical properties of the sludge have a

great influence on the selection of the sewage sludge treat­

ment procedure.

Activated sludge arises spontaneously in an activated sludge

installation from the micro-organisms present in the waste

water. It consists of flocculated bacteria, protozoa, etc.

and adsorbed material.

The chemical composition is about

C118H170051N17p

The bacteria have a diameter of 0.5 to 1.5µ and are seldom

longer than 10µ.

Their slimy skin causes them to be grouped together into tenu­

ous flocks which may have characteristic dimensions of 20 to

100µ.

The concentration of organic contaminations in sewage and sew­

age sludge is often characterised by the biological oxygen de­

mand (BOD). This is the amount of oxygen which is consumed per

unit volume by the action of micro-organisms on the contamina­

tions (77). In general the BOD is expressed in mg/1.

Another characterisation of the concentration which will be

used in this thesis is the chemical oxygen demand (COD).

The COD is the amount of oxygen necessary per unit volume for

oxidation with a dichromate-sulfuric acid mixture under stand­

ard conditions (77). In general the COD is also expressed in

mg/1. However, in this thesis practical units are used, so the

COD is expressed in kg/m3.

Since the break-down by the dichromate-sulfuric acid mixture

in general proceeds further than by the action of the micro­

organisms, the COD is higher than the BOD. The relation be-

4

tween BOD and COD will frequently vary.For domestic waste wat­

er Hunter <i27) reports that

COD ::::: 2 BOD

1.4. SEWAGE SLUDGE TREATMENT AND DISPOSAL

The numerous sludge treatment procedures may be grouped into

four major categories, indicated in figure 1.2.

raw sludge

,--concentration

digestion

dewatering

combustion

final disposal

Figure 1.2.

Basic steps in sewage sZudge treatment.

Which kind of combination of these procedures provides the

most economic and reliable solut~on, depends on the nature and concentration of the sludge, on the selected disposal method and on local situations.

The concentration step of the sludge suspension, indicated in

figure 1.2, yields a more economic treatment in subsequent p:ro­cesses ( 5, 14) • Anaerobic sludge digestion. involves biological breakdown of organic matter in the absence of oxygen by anaerobic bacte-

ria.

5

Dewatering by filtration, centrifuging or by the use of drying

beds, results in a sludge of a more or less solid state. How­

ever, flocculation agents have to be added (15, 16, 17, 18).

Increase of filtration rate can also be obtained by heating

the sludge for half an hour at 180-200°c and 10-15 atmospheres,

resulting in the disappearance of the colloidal structure (13,

21, 22, 23, 119).

This heat treatment process produces a sludge with an offen­

sive smell, which, however, disappears when air is also fed to

the reactor, resulting in partial oxidation (2, 4, 15).

Combustion reduces the volume of the solids. The final product

is a mineral and odourless ash.

Two groups of combustion processes can be distinguished:

(i) without previous dewatering (20, 26, 27, 28, 38),

e.g. the wet-air oxidation process;

(ii) with previous dewatering and heat drying (20, 29, 30),

e.g. combustion in a fluidised bed.

The wet-air oxidation process is a new and promising process

by which the organic matter in the sludge is oxidised in the

liquid water phase at elevated temperatures and pressures

(220-300°c, 60-125 atmospheres ) in the presence of air (27,

28, 38)· The process is diagrammatically represented in figure

1.3.

In the heat exchanger, sludge and air are preheated prior to

being admitted to the reactor. In the gas-bubble slurry reac­

tor exothermic oxidation of the sludge takes place by which

the temperature rises about 30°c. The heat content of the ef­

fluent of the reactor is used to preheat the sludge and air

feed. After heat exchange, gas and liquid phases are separated

and expanded to atmospheric pressure. For a large plant reduc­

tion in compression energy cost can be accomplished by using

an expansion engine for the separated gases.

In selecting the method for the final disposal of the treated

sludge, due consideration should be given to the requirement

6

sludge pump

sludge

Figure 1.3.

eactor

expansion engin

stack gas

air

compressor

oxidised sludge

Wet-air oxidation process.

that public heal th or safety· shall not be impaired and no new

pollution problem is generated.

When modern treatment procedures like wet-air oxidation are

applied, the final disposal of the sterile inorganic ashes

will give no problems.

7

Chapter 2

0 l

-

THE WET-AIR OXIDATION PROCESS

2.1. EVOLUTION OF THE PROCESS

The history of wet-air oxidation starts in 1912 when

Strehlenert (41) patents a method for the treatment of spent

sulfite liquor from paper mills with compressed air at 180°c.

In later versions of the process the oxidation of the paper

mill effluent is performed at temperatures ranging from 230 -

330°c. This method was first patented in Sweden in 1949 by

Cederquist (42, 44, 45); his process has not been applied in

practice (46).

Independently, Zinunermann patented nearly the same process(43)

in the U.S.A. in 1950. This patent was followed by many others

for a diversity of operating conditions and process perform­

ances (e.g. 47/69). The development and promotion of his pro­

cess was carried out by the ZIMPRO (ZIMmermann PROcess) divi­

sion of Sterling Drug. The first conunercial installation for

the oxidation of spent sulfite liquor, had to be shut down

8

after a short time because of corrosion (70, 71, 72). Till

February 1969, 18 other installations were sold (73). The

largest installation is located near Chicago and has a capac­

ity of 200 tons of dry sludge per day.

A patent for a system of wet-air oxidation in a deep shaft ex­

tending into the earth was awarded to Bauer (112). The depth

of the shaft was made sufficient to provide the required high

pressure.

The first patents claimed nearly complete oxidation and atten­

tion was focussed on paper mill effluent. Attention was changed to sewage sludge and the advised degree of oxidation was gradually reduced. Nowadays 5-20% oxidation

is applied at relatively low temperatures of 180 - 200°c, with the prime object of obtaining a better drainable sludge so

that the original combustion process is transformed into a conditioning process prior to dewatering.

However, owing to the attractive possibilities of the orig·inal combustion version and owing to the fact that this process has

never been approached in a process-engineering manner, we car­ried out an investigation of the high temperature version,

which is embodied in this thesis.

2.2. OXIDATION "ROUTES"

Sewage sludge consists of solid particles suspended in waste

water. In the wet-air oxidation process the oxygen is supplied

as gaseous air; therefore, a three-phase system is provided in

the reactor. The oxygen diffuses from the gas-bubbles through the gas-liquid interface into the suspension, where it reacts with the solid sludge particles.

At the elevated temperatures the organic polymeric structures

of the sludge are hydrolised to smaller soluble molecules (23, 120).

9

Teletzke (75) observed the formation of free amino acids, free

fatty acids and lower sugar molecules.

Hurwitz ( 76 ) presents data from which it follows that the

fraction of organic material which is dissolved increases rap­

idly with temperature and approaches 1 at 260°c.

As a result of this dissolution the oxidation of sewage sludge

can proceed via the solid particles as well as via the dis­

solved matter.

Takamatsu (129) assumed that the oxidation only proceeds

through dissolved matter, but he did not prove this experimen­

tally.

In this respect it can be referred to a patent of a system of

wet-air oxidation in which the sludge is first hydrolised as

far as possible. After settling of the solid residue which

takes place under the high temperature and pressure, only the

sludge solution is oxidised (114).

By the wet-air oxidation volatile products like acetic acid

might be generated. Oxidation of volatile products only takes

place in the liquid; in the gas phase no oxidation was ever

observed at the applied conditions (70,74).

2.3. INFLUENCE OF PROCESS PARAMETERS

In fig. 2.1 the influence of the temperature and the reaction

time on the COD reduction of a sludge suspension is repre­

sented. The data have been taken from Zimmermann (70). Total

pressure,

mentioned.

starting concentration and kind of sludge were not

Nevertheless it follows from this graph that com-

plete

3oo0 c. fast

oxidation is not possible at temperatures lower than

It is also seen that there is at first a relatively

oxidation, which is followed ny a very slow reaction.

The COD reduction reached after the relatively fast reaction

can be considered as a maximum conversion. As follows from

figure 2.1, this maximum is a function of temperature. Nearly

all literature only deals with maximum conversion.

10

ioor-~-:::::::==================:---i

80 ~OD

reduction60

(%)

40

20

Figure 2.1.

time (h)

3oo0 c 250

200

150

100

InfZuenae of temperature and time on COD reduation.

Figure 2.2 shows nearly all published data. Temperature: and

time are indicated in the graph. The other conditions are pre­

sented in the tables 2.1 and 2.2.

The region of reaction times of 60 to 120 minutes represents

batch experiments with activated,primary and digested sludges.

The curve (t = 00 ) indicates that after "maximum" conversion is

attained still a considerable amount of COD reduction can be

reached if the reaction is given enough time.

In the first publications of Zimpro (70,74) it was stated that

the only function of the high pressure was to reduce evapora­

tion in the reactor, which is necessary since the reaction

stops if no liquid water phase is present. However, Hurwitz

(79) found that higher pressures also increased the capacity

of an installation. The influence of the oxygen partial pres-

sure was suggested by the experiments of Abel (80) and of

Pepelyaev (81).

In the literature no investigation about the influence of the

sludge concentration on the conversion rate could be found.

11

80

60 COD

rec11.lction ( % )

40

20

c

Tab le 2. 1.

temperature (0 c)

Figure 2. 2.

Survey of data from literature.

(): aommeroial installation

see also tables 2.1 and 2.2.

Figures in graph are residenae times in min.

Some aommeroial installations represented in fig. 2.2.

location kind of starting temp pressure COD re-

sludge COD (kg/m3 ) (OC) (atm) duction(%)

a Chicago activated 48 270 125 80 b ·Wausau primary 62 260 115 75

c Wheeling primary 43-95 260 83 65 d Blind primary 41-97 235 53 79

Brook

e Milwaukee raw/ 36-66 200 34 70 digested primary

12

ref.

82,90

90

88,90

89,90

90,91

Tabie 2.2.

Experimentai conditions of experiments represented in fig.2.2.

residence kind of COD temp. pressure batch/ ref.

time (min) sludge 0 3

( OC ) (atm) contin. (kg/m )

"' ? ? till 200 ? batch 68,69

180 activated 51-75 260 83 cont. 79

60 primary, 15-100 100-300 ? batch 78 activated

60 primary 62,58, 100-300 ? batch 76 activated 31 digested

60 primary 58 150'-250 ? batch 75

38 activated 48-54 277 122 cont. 79

35 activated 40-49 274 102 cont. 79

20 ? ? 150-225 ? cont. 68

19 activated 28-96 243-272 ? cont. 65 activ/prim.

6 primary 41-107 200-250 32-68 cont. 68

2.4. END PRODUCTS

For the final disposal of the solid end products, these have

·to be separated from the liquid effluent. This process has

been studied by Walters and Ettelt (83) and by Teletzke (84,

85) •

The liquid effluent may contain a considerable amount of dis­

solved matter and could be returned to the biological oxi­

dation unit.

The chemical composition

Teletzke (75).

2 • 5 • COST OF THE PROCESS

of the effluent has been studied by .

The cost of the high temperature version of the wet-air oxi­

dation process is determined by the desired COD reduction.

Teletzke (24, 25) gives the relative cost as a function of COD

reduction.

13

The absolute cost presented in the literature shows a consid­

erable scatter. For the Chicago installation Goldstein (82)

reports $23 per ton dry sludge, which does not include inter­

est on capital investment. Five years later, Dalton (32) mentions $50 per ton for the

same installation. He also refers to cost of other sludge

treatment processes at the Chicago sewage plant.

2.6. LITERATURE REVIEWS

In the literature a great many reviews on wet-air oxidation

have appeared in several languages. In English, e.g. refs.

(92)/(97), (111); in German, e.g. refs. (98)/(102); in Polish ,

e.g. refs. (103, 104); in Dutch, ref. (105); in Swedish,ref.

(106). Some additional experimental data can be found in refs.

(107, 108, 109).

2.7. CONCLUSIONS

From the literature it follows that the wet-air oxidation of

sludge is influenced by the temperature

Quantitative influences are not published.

and the pressure.

Quite clear is the

effect of temperature on the maximum conversion; however, the

rate by which this is achieved is unknown. Furthermore, no in­formation was found concerning

(i) The kinetics of the reaction.

(ii) The influence of oxygen transfer from the gas phase into

the suspension.

(iii) The dissolution of solid sludge particles,which may have

an effect on the oxidation rate by, e.g., different sta­bilities of solid and hydrolised, dissolved sludge.

(iv) The influence of mixing in the reactor.

Without knowledge of or insight into these factors, a proper

design of a wet-air oxidation process cannot be expected. Only

after realisation of such a design can a fair comparison of

cost with conventional sludge treatment procedures be made.

14

Chapter 3 THE OXIDA1.ION OF A GLUCOSE

SOLUTION AS A MODEL SLUDGE

In order to obtain insight into the general behaviour of wet­air oxidations, the research project was started with the oxi­

dation in a two-phase gas-liquid system of a model sludge for

which a solution of glucose was selected, glucose being repre­

sentative of the group of carbohydrates.

3.1. PRELIMINARY EXPERIMENTS

Semi-batch experiments were carried out in a one-litre elec­

trically heated autoclave which is diagrammatically shown in

figure 3.1. For each experiment about half a litre of water

was heated in the reactor to the desired reaction temperature.

A vertically moving agitator (shaker), placed inside the reac-

15

oxygen/air spent gas

injector cooler

t

sample

cooler

Figure 3.1.

Semi-batoh installation.

tor, provided mixing of the contents of the reactor. 20 to

40 ml of concentrated glucose solution was injected and air or

oxygen was fed into the system and passed over or through the

glucose solution. The spent gases were cooled in a condenser

from which the condensed steam flowed back into the reactor. A

dip-pipe enabled samples of the liquid in the reactor to be

drawn.

In order to ensure that the time necessary for mixing the liq­

uid phase homogeneously was short enough not to influence the

outcome of the experiments,a conductivity electrode was placed

in the reactor. Then a concentrated sodium chloride solution

was injected and it was found that the solution was nearly

completely mixed up in five seconds at the shaker frequency of

144 min- 1 .

By heating a glucose solution a number of sequential dehydra­

tion and polymerisation reactions take place. By these thermal

reactions the solution is coloured more or less brown, and ob­

tains a sweet smell of caramel (46).These reactions could also

16

take place during oxidation experiments and oxidation could

proceed directly from glucose as well as via thermal reaction

products.

The influence of thermal treatment on subsequent oxidation was

studied by changing the time between injection of glucose and

supply of oxygen 1.which changes the extent of thermal reaction.

The results of these experiments are presented in figure 3.2.

Because of the large number of possible reaction products, it

was decided not to follow each component individually, .but to

trace the total amount of oxidisable material. The concentra­

tion of this oxidisable material is expressed in the chemical

oxygen demand (COD). The experiments were carried out at 178 and 200°c and 50 atm.

At the temperature of 178°c air was passed through the solu­

tion at. a flow rate of 1 Nm3 /h. · The reactor was filled with

15

COD

(kg/m3)

10

5

0

(1) non-preheated, 178°c

(2) preheated, 178°c · '-------...... <

' -......, (3) preheated, 200°c \ ........ , \ ' ', ( 4) non-preheated, 200°c

....... , ....... ,

' ...................

............ (1) (3) .... _ -----

(4)

10 2

time (min)

Figu'I'e 3.2.

Effeat of p'I'eheating on subsequent o::cidation of gluaose.

17

500 ml water enabling the shaker to force gas bubbles into the

solution from the gas space above it (see 3.1.3.). The total -1 pressure was always 50 atm and the shaker frequency 144 min •

In the experiment belonging to curve (1), the temperature was

178°c and air was admitted to the reactor immediately after

the injection of glucose. The curve indicates that it takes

several minutes before noticeable oxidation takes place. Ap­

parently, active material has first to be formed.

The experiment of curve (2) was carried out at the same tem­

perature, but now the glucose solution was heated for 10 mi­

nutes in the absence of air so that only thermal reaction *)

took place. Then air was admitted and now, as follows from

curve (2), the oxidation started nearly immediately.

This could be understood by assuming that the active material

is formed by thermal reaction. The active material could be a

catalyst for the oxidation of glucose and other thermal reac­

tion products, but it is also possible that we are dealing

with a consecutive reaction, which would mean that thermal re­

action products are more reactive than glucose. De Wilt (122)

has shown that the oxidation of glucose in alkaline solutions

at about 60°c proceeds through enolate ions, generated from

glucose by thermal reactions, which makes the last mechanism

most likely to occur in our case.

The experiment at 178°c with non-preheated glucose was re­peated with pure oxygen passed over the solu.tion at 50 atm, a

shaker frequency of 144 min- 1 ,and a gas flow rate of 0.6 Nm3/h.

The reactor was now filled with more than 650 ml water, so

that no bubbles were forced into the solution (see 3.1.3.),

resulting in a much lower gas-liquid interfacial area compared

with the above described experiments. The curve obtained coin­

cides nearly completely with curve (1). So, also with pure ox­

ygen, a "lag-phase" is obtained.

*) By thermal reaction we mean those reactions which take

place without the influence of oxygen.

18

Curve (3) represents the experiment with non-preheated glucose

at 200°c. Oxygen was passed over the solution at a flow rate

of 0.6 Nm3/h, the reactor being filled with more than 650 ml

water. At this temperature oxidation seems to start almost

immediately. By preheating at this temperature for 10 minutes,

the subsequent oxidation, indicated by curve (4), proceeds

more slowly compared with the non-preheated glucose.

At 200°c thermal reactions will have proceeded further than at

178°c, so, apparently some thermal reaction is necessary for

rapid oxidation; however, too much thermal reaction decreases

the rate of oxidation. This could be explained by assuming

that the generated active material is degraded by further

thermal reaction.

When the solution was preheated for over one hour at 200°c,

almost all glucose was converted into a carbon-like product

which was deposited on the reactor wall and the shaker.

From figure 3.2 it follows also that complete oxidation is not

reached at the selected temperatures, but a maximum conversion

is obtained. Since only a slight difference occurred in

maximum conversion for a preheated glucose solution and a non-­

preheated one, this maximum conversion can only partly be in­

fluenced by thermal reaction products. Apparently, it is

mainly determined by oxidation products, which resist further

oxidation under the experimental conditions.

In order to get an impression of the influence of the shaker

frequency on the gas dispersion, model experiments at room

temperature were carried out in a glass vessel of the same di­

mensions as the autoclave, the glucose solution being replaced

by hexane, which at room temperature has a viscosity and sur­

face tension near those of water at 200°c.

It was observed that the gas dispersion was strongly inf lu-

19

enced by the liquid level in the reactor. At high levels, when

the upper blade of the shaker was completely submerged in the

liquid as well as the gas-inlet, only a very small fractional

gas hold-up (less than 0.01) was observed, which originated

from the bubbles passing through the liquid.

At lower liquid levels, when the shaker passed through the

gas-liquid interface, a large amount of gas phase was forced

into the solution as bubbles resulting in a gas hold-up of a­

bout 0.11 at a shaker frequency of 144 min- 1 • At low frequen­

cies all bubbles escaped from the liquid between two shaker

cycles. The observed average bubble diameter {db) seemed to be

independent of the shaker frequency and was 2 to 3 mm.However,

the average number of bubbles in the liquid increased with the

shaker frequency, resulting in an increase of the average gas--1 liquid interface. At the maximum frequency of 144 min and a

liquid volume of 600 ml, the fractional gas hold-up {£) was

0.11. Consequently the specific surface area of the bubbles,

6£ ab = d'

b

was about 270 m- 1

When changing the liquid level, it was observed that for liq­

uid hold-ups larger than about 650 ml the shaker was complete­

ly submerged during shaking.

So, by applying a liquid hold-up exceeding 650 ml a relatively

low specific surface area must be expected.

The influence of the shaker frequency on the oxidation rate

was determined also at 200°c and 50 atm with an air flow of l

Nm3/h. The liquid hold-up was 500 ml so that according to the

model experiments it must be expected that air had been forced

into the solution. The results are shown in figure 3.3.

It follows from this graph that the shaker frequency influ­

ences the conversion rate. This indicates that the latter de­

pends on the gas-liquid interface, which means that in these

20

20 0 20 min-l

c:l 35 II

+ 60 II

15 x 144 II

( COD

(kg/m3)

10

5

0 0 1 1 2

time (min)

Figure 3.3.

Effeat of shaker frequency on o~idation at 200°c and 50 atm.

experiments the conversion rate is also determined by dif fu­sion.

From the experiments discussed in 3.1.3. it followed that the

oxidation rate of glucose under the conditions of the experi­

ments was lowered by diffusion limitation. so, a direct deter­

mination of the kinetics of the reaction was not possible with

such experiments. owing to the relatively high solubility of

oxygen above l00°c (see fig. 3.14), the oxidation of glucose

could be carried out in a homogeneous system. The results are

shown in figure 3.4.

21

0.6·

COD (kg/m3)

0.2

0

Figure J.4.

200°c

[ob] = 1 kg/m 3

50

time (sec)

Homogeneous o~idation of gZuaose.

In the autoclave 500 ml water of 200°c was saturated with oxy­gen to a concentration of 1 kg/m3• The shaker frequency was 144 min-1 • Then some glucose solution was injected, resulting in a initial COD of 0.6 kg/m3 • With regard to complete oxi­dation, a surplus of oxyg.en of nearly 70% was present. Although the oxygen consumed by the reaction will be supplied by mass transfer from the gas phase, the original surplus of

oxygen present in the liquid will guarantee that diffusion limitation has only little influence on the observed reaction rate.

From figure 3.4 it may be seen that the homogeneous oxidation

proceeded so rapidly that after 30 seconds, when the first sample was taken, the maximum conversion was already nearly reached. For the heterogeneous oxidations which were shown in figure 3.3, it took 10 minutes to reach the same degree of conversion, while the equilibrium oxygen concentration there was 1. 4 kg/m3• It is quite clear that homogeneous experiments carried out in the way described here, are too fast for studying the kinetics of the oxidation of glucose.

22

As mentioned in 3.1.2. complete oxidation was not reached. The

maximum conversion, defined as the COD reduction after one

hour, was determined as a function of temperature.

Figure 3.5 shows that it increases from 75% at 170°c to 90% at 26o0 c. Results are also presented which were obtained from a

further oxidation of samples of continuous flow experiments

(see 3.2.). As follows from the graph, the maximum conversion

data are in good agreement with the results of Yunis (71), who

studied the wet-air oxidation of glucose in the presence of

hydrogen peroxide and ferric ions. This indicates that prod­

ucts resisting wet-air oxidation also resist the action of hy­drogen peroxide/ferric ions.

100

maximum COD

reduction (%)

75

50

25

FiguPe 3.5.

T glucose

• preoxidised glue.

• glucose+ H20 2 (71)

• sugar (70)

temperature (0 c)

Ma~imum COD Peduation of gtuaose and sugaP.

23

From the preceeding sections it followed that complete oxida-o tion was not reached even at a temperature of 260 c. It was

found from the continuous flow experiments that the amount of

coo that could be removed, the "effective" coo [c], is a pa­

rameter by which the overall conversion rate per unit volume

can be described with a half order in effective coo. For the

semi-batch experiments this results in

- d£~l =constant* [c]~

Integration of eq. (3.1), using the boundary condition

t = O, [c] = [c0],

results in

[c ]~ - [c]~ =constant* t 0

By introducing y = 1s1 this transforms into [co]

1 - YJ..z = constant *t

[ c ] J..z 0

(3.1)

(3.2)

(3.3)

Since this description does not include generation of the ac-

·tive matter, it may only be applied after the "lag-phase".

In figure 3.6 the results of the experiments discussed in

3.1.2. are plotted as 1 - y~ against time. This results in

straight lines up to 80% of the maximum conversion.

The deviation at higher conversions might be the result of the

formation of rather stable oxidation products.

Figure 3.6 also shows that when preheating for 10 minutes in

the absence of oxygen at 200°c, the subsequent oxidation · pro­

ceeds more slowly compared with a non-preheated solution.

Finally it follows from this graph that 10 minutes preheating

at 178°c results in a higher conversion rate compared with

non-preheated glucose.

24

1. 00

l-y15

(1) non-preheated (2) preheated

time (min) 178°c

Figure 3.6.

( 1)

(2)

non-preheated (2) preheated

10 time

0(min)

200 c

Results plotted as half-order aonversion rate.

3.2. CONTINUOUS FLOW EXPERIMENTS

1. 00

0.75

a.so

0.25

o.oo

Figure 3.7 shows the flow sheet of the bench-scale continuous flow installation.

The reactor consisted of a stainless steel bubble column with an internal diameter of 0.04m and a length of 1.00 m.

Glucose solution and air

steel electric preheaters

respectively, and fed into

heater : 12 1).

were heated separately in stainless

with capacities of 7.5 and 0.5 kW

the reactor (volume of liquid pre-

25

sludge

Figure 3.?.

Continuous fto~ instaiiation.

i--.,_~pen t gas

air

A: studge pump; B: preheater; C: reaator; D: aooier;

E: separator; F: ao~pressor.

Air was dispersed by a multiple orifice distributor, diagram­matically shown in figure 3.8, and rose in co-current with the liquid through the reactor.In some experiments the air was fed

into the system just before the sludge preheater so that glu-

26

sludge inlet

air inlet

Figure 3.8.

Injeation system.

cose and air were heated simultaneously. This was not the nor­

mal procedure, however, as in that case the locus of oxidation

was not clearly defined.

The gas and liquid phases were removed at the top of the reac­

tor and then cooled rapidly. After pressure expansion, gas and liquid were separated. The temperature in the reactor was

measured by thermocouples in cylindrical wells at the bottom, half-way, and at the top of the reactor. The temperature was

controlled by the preheaters and could be kept constant within ' 0 1 - 2 c. The total pressure was measured at the top of the re-

actor and was kept constant by a back-pressure valve. Liquid

phase samples could be drawn half-way and at the bottom of the reactor. The oxygen concentration in the spent gas was meas­ured with an oxygen analyser. In order to change the oxygen pressure, oxygen or nitrogen could be mixed with the air feed

of the compressor. The experiments were all carried out at a total pressure of 50

atm and a gas flow rate of 1.7 Nm3/h. Unless mentioned other­wise, air was used while the liquid feed rate was l0- 2m3/h,the

feed having a COD of 35 kg/m3 • owing to variations in thee­vaporation with temperature, the average residence time was

dependent on the temperature.

In figure 3.9 curve (1) represents the influence of the tem­

perature in the bubble column on the COD reduction. At a temperature of about 2oa0 c the COD reduction suddenly dropped. Presumably this was caused by a too far proceeded thermal reaction in the preheater. In the semi-batch experi­

ments it was already shown that some thermal reaction was nec­essary for fast oxidation, but that too much thermal reaction

decreased the oxidation rate (see 3.1.2.). In accordance with the foregoing, the oxidation will proceed

faster when it takes place during or after the first steps of

the thermal reactions, which means in the sludge preheater. In

27

100

0 air straight into reactor + air via sludge preheater

75 0

COD ~+- (2)

reduction

.~: '.'.!-,.........+

(%)

50 0

I b I A" (1)

I --.__.

25 oh

0 190 210 230

temperature (0 c) Figure 3. 9.

Effect of temperature on continuous ftow oxidation of gtucose.

order to test this hypothesis, air was mixed with the glucose

solution before it entered the sludge preheater so that glu­

cose and air were heated simultaneously.

Also with this arrangement the influence of temperature was

determined. The results are given in figure 3.9 by curve (2)

which shows that the drop in conversion has disappeared.

Up to

though

2os0 c the COD reduction is

the gas-liquid contact time

somewhat lower now, even

is about twice as high as

for curve (1). No gas distributor was used, however, which

definitely results in a lower interfacial area than in the ex­

periments described above. This again confirms that the oxi­

dation rate is dependent on the gas-liquid interface, as con­

cluded from the semi-batch experiments.

Since the locus of oxidation is only defined when no oxidation

takes place in the preheater, the experiments discussed in the

28

following were always carried out in such a way that air and

glucose were heated separately.

3.2.3. ~n!1Y~n£~_g£_£QQ~22n2~n~=2s!2n_!n_sh~-£~~§_2n§_g~yg~u

E=~!!Y=~

At a temperature of 220°c the influence of the COD concentra­

tion in the feed was measured. Results are presented in table 3.1. It follows from these data that the conversion rate in­

creases only slightly with the COD concentration,while the COD reduction drops. This indicates that the conversion order in the organic material will be somewhere between zero and one.

TabZe 3.1.

InfZuenoe of COD in the feed.

CODf COD reduction conversion rate (kg/m3 ) (%) (kg/h)

20.3 72 0.146

29.4 60 0.170

33.9 55 0.186

The oxygen pressure

with the air feed

was changed by mixing oxygen or nitrogen

of the compressor, keeping total pressure

and gas flow constant.

The influence of the oxygen pressure was examined at several

reaction conditions. The results are presented in figure 3.10.

At the temperatures of 207 and 213°c the liquid flow rate was 5.2 x 10-3 m3/h, while the COD concentration in the feed was

3 . -2 3 38 kg/m .At the other temperatures these values were 10 m /h and 37 kg/m3 , respectively.

From figure 3.10 it can be concluded that the oxygen pressure

affects the conversion rate as could be expected, but that this effect levels out at higher pressures.

29

0. 15 .

conversion rate

(kg/h)

0. IO ·

o.os

0

Figure 3.10. oxygen pressure (ata)

Effect of oxygen pressure on continuous flow oxidation of glucose.

Regarding the influence of concentrations, it was found that -the oxygen flux through the interface q

0 can be expressed in

terms of oxygen pressure, p, and effective COD, [c]

constant * [ c] ?..i (3.4)

3.3. THEORETICAL ANALYSIS OF RESULTS AND DISCUSSION

For a proper analysis of the results, the mixing state in the

reactor must be known. Therefore, residence time distributions

of the liquid phase were measured, using sodium chloride as a

tracer.

30

It was possible to describe the measured distributions as plug-flow with axial mixing (84, 85). The corresponding Pc§clet numbers were of the order of 3.5, showing that only a few mix­ing stages were present. For a definition of the PAclet number see 3.3.2. Figure 3.11 shows a measured and a calculated resi­dence time distribution.

c*

o.s-

~~ measurements calculated for Pe = 3. 5

1 2 0 --~~~~--~~~~..__~~~---'

t/t

Figure 5.11.

Measured and aaZauZated residenae time distribution.

In experiments at 190°c, samples were taken from the liquid in the reactor, in which the nitrogen concentration, [cN ], was

2 measured.

The equilibrium concent.ration belonging to the nitrogen pres-

* sure in the reactor [cN ] was calculated, using Pray's (78) solubility data for 2 pure water. From these two values the overall volumetric mass transfer co­efficient, K1a, was determined, using the equation

in which ~t is the liquid flow rate.

31

A mean value for Kia of 4.1 x 10-2 sec- 1 was found. Using this

value for the physical absorption rate of oxygen, and assuming

zero oxygen concentration in the liquid, the maximum physical

absorption rate of oxygen was calculated to be 50 g/h.

Since the observed conversion rate was 145 g/h, the oxidation must take place mainly within the diffusion layer around the

gas bubbles.

3.3.2. Model for the macro kinetics ----------------------------From the continuous flow experiments it followed that the ab­sorption rate of oxygen is proportional to the square root of

both oxygen pressure and effective coo. In the theoretical a­

nalysis presented in 3.3.3. an expression will be derived for

the constant factor in equation (3.4). Introduced in the ab­sorption rate per unit interfacial area, q

0, becomes

= v 2D0~[c]p (3.5)

With this expression a mathematical model of the conversion

rate obtained in a continuous flow reactor was set up, includ­

ing combined mass transfer and reaction, convection and mix­

ing. The model is based upon the following starting-points:

(i) Equation (3.5) describes the local oxygen transfer rate

through the gas-liquid interface, using the local values of [c] and p.

{ii) The residence time distribution in the liquid phase can be described as plug flow with axial mixing.

(iii) The gas flows in co-current with the liquid in pure plug­

flow.

{iv) The radial mixing is so high that in radial direction

the concentration profiles are flat.

{v) The temperature is uniform throughout the reactor.

32

(vi) The gas flow, liquid flow and interfacial area are uni­

form throughout the reactor, which means that all evap­

oration takes place at the inlet.

(vii) The reactor is operated in steady state.

A mass balance of COD over the liquid phase between the cross­

sections in the reactor at the heights x and x + ~x results in

the following differential equation

d r,..l d2 [C] L L ~ - u .::..i..£..i.. + E - - - p~[c]~ a= 0 JI, dx dx2

(3.Sa)

where

UR, . superficial liquid velocity .

x . length co-ordinate in reactor . E : eddy diffusivity taken per unit reactor volume

a • specific gas-liquid interface taken per unit reactor . volume

A mass balance of oxygen over the gas phase results in

u f!l_o k ....S: ~ + p~[c]~ --St: a = 0 RT dx H (3.6)

where

R gas constant

T absolute temperature

u superficial gas velocity. g

The boundary conditions for the two simultaneous differential

equations are

x = 0 p = pf

[cf] [co] E (fil.£1) - =

UR, dx x = 0

x = x -aJ~J = 0

33

In this the subscript f refers to feed conditions. By intro­

ducing the following dimensionless variables and parameters

* [c]/[cf] y = 1T = p/pf

a = x/X

Pe UR,X

(Peclet nwnber) = E 2D kRT

Nr o- a X(number of conversion stages) = H U.v,Ug

M uR.[cf]RT feed rate of oxidisable material = = ,

ugpf feed rate of oxygen

the equations (3.Sa) and (3.6) can be transformed into the di­mensionless equations

dy* + L d2~* - ~1T~YR~ = 0 (3.7) do Pe do M

d1T + N M~yR~ 1T ~ = 0 ( 3. 8) da r

while the boundary conditions become

o = 0 1T = 1

l * L dy., - y = Pe do

o = 1 dy* do = 0

The two simultaneous equations were solved on an analogue com­

puter. Some of the results are given in the figures 3.12 and

3.13. These and additional results have already been published

elsewhere (125). In order to test the model, in figure 3.12 the results of ex­

periments presented in 3.2.3. have also been included. For each series of these experiments Nr and Pe were constant,

34

100

75

conversion

(%)

50

25

Figure 3. 12.

Effeat of M and N NP is indicated i~

+ 220°C

• 213

y 207

• 186

on aonvePsion of gZuaose foP Pe= 3.5. the gPaph.

while the influence of M was determined by changing the COD

of the feed solution or the oxygen pressure of the feed gas.

The conversion is based upon the maximum conversion and is,

therefore, expressed as a percentage of the reduction of the

effective COD.

In the graph the line for complete utilisation of the oxygen

feed is also drawn, represented by the equation

conversion = lOO (%) M

It follows from figure 3.12 that there is a reasonably good

agreement between the calculated curves and the measurements.

Another comparison is represented in figure 3.13. This graph

shows several measured local concentrations in the reactor and

the corresponding calculated concentration profile. This also

shows an acceptable agreement.

35

1.0

y* 0 data points

calculated curve

for Pe = 3.5

0"" Nr = 1. 25 0.5

M = 0.5 0

0 0 0.5 1.0

a Figure 3. 13.

Calculated and measured concentration distribution in reactor.

It may be concluded that the mathematical model gives a fair

description of the phenomena in the reactor which influence

the overall conversion.

3.3.3. Model for the micro kinetics ----------------------------To understand the square root dependence of the absorption

rate on the concentrations, models for mass transfer with

chemical reaction were examined. From the experiments it fol­

lowed that the reaction was fast compared with diffusion of

oxygen and, therefore, that the oxidation took place in a thin

film near the gas liquid interface.

In this section it will be demonstrated that a penetration

model of absorption of oxygen, followed by rapid oxidation,

which chemical-kinetically is of first order in organic mate­

rial and zero order in oxygen, results in the observed conver­

sion orders.

36

The physical picture on which this penetration model is based

is that at a time zero a liquid element is contacted with an

air bubble. During a time T penetration of oxygen and oxi­

dation takes place. Then the element is replaced by another

one and the process starts again. The absorption of oxygen in

the element followed by the chemical reaction is mathemati­

cally represented by

( 3. 9)

where D0

is the diffusivity of oxygen in the liquid, [oJ the

oxygen concentration and B the reaction term.

For a reaction which is zero order in oxygen and first order

in reactant present in the liquid, B is given by

B = ~[c] (3.10)

where ~ is the first order reaction rate constant and [c] the

concentration of the reactant. The concentration [c] is de­

scribed by the equation

= llil at + R (3.11)

in which Dr is the diffusivity of the reactant in the liquid.

At time zero the initial concentration of organic material is

uniformly equivalent to [ch]' the concentration of reactant in

the bulk of the .liquid. Because of reaction, reduction of the

concentration of organic material will occur which will be

most pronounced at the interface,since the reaction penetrates

from the interface inwards. In the lapse of time T the quan­

tity of organic matter which has disappeared per unit volume

f the liquid in the neighbourhood of the surf ace is smaller

than ~T [ch]. The quantity present at time zero is [ch]. A

sufficient condition that the variation of [c] may be ignored

during lapse of time T is that

37

or

kT << 1 (3.12)

This condition is unnecessarilly stringent since suppletion by diffusion is neglected.

Particularly this is the case when

D (cb]>>D [o.] r o J.

or

Because of the relatively low solubility of oxygen (78) the

left hand term is of the order of 20, which means that condi­

tion (3.12) definitely is too stringent and can be transferred

into

1sT < 1. (3.13)

If condition(3.13) is fulfilled [c] may be considered constant

during time Tat the value [cb]. By this the reaction term in

equation (3.9) is a constant. An analogous differential equa­

tion was solved numerically by Astarita (90).A stationary con­

centration profile of oxygen will eventually be obtained which

is nearly reached when

(3.14)

If this condition is fulfilled, the time derivative in equa­

tion (3.9) can be neglected, and if, in addition, (3.13) is

also fulfilled, equations (3.9) and (3.10) are reduced to

38

[c] = constant = [cb]

2 Do d [o] = k[c ]

dz2 - b •

.(3.15)

(3.16)

The steady state concentration profile is extending between

the interface and a distance o where [o] = O. So, the boundary

conditions are

z = 0 [o] = constant = [oi]

z = o : [o] = Q,

The value of o can be obtained from a mass balance. Because of the steady state the amount of oxygen passing through the in­terface per unit time equals the amount of oxygen which dis­appears by oxidation in the layer between z = o and z = o • So,

- D (fil2.l) o dz z = o

The solution of (3.16) with the boundary conditions is

2 19..L = 1 - r; 12+ L [o.] 2

1

in which

r; = z

(3.17)

(3.18)

From this it follows that the steady state absorption rate of oxygen per unit interfacial area, qo,s' equals

fil.21 -qo,s = - D v 20 o!s [Cb] [ O i] • 0 ( dz )z=o - (3.19)

If the gas-phase resistance can be neglected, [oi] is related to the oxygen pressure p by

(3.20)

39

where H is the Henry coefficient. For pure water at elevated

temperatures and pressures Pray has shown that the Henry coef­

ficient is only dependent on temperature (78).

Assuming that the organic material does not influence the

Henry coefficient, equation (3.23) can be transformed into

(3.21)

This is the steady state absorption rate for e + 00 • In general

the average absorption rate during the time of contact is

given by

and depends on e. From Astarita's data (90) q0

was calculated.

The results are shown in figure 3.15 where the per cent devia­

tion of q0

from the steady state absorption rate q0

,s is given

as a function of e. It follows from the graph that even when a

Steady State COnCentratiOn profile iS established (e~o.4) I the

average absorption rate is still 40% above the steady state absorption rate which is given by equation(3.19) .If we allow a

deviation of 10%,the criterion for a steady absorption rate is

e ~ 2. (3.22)

This combined with condition (3.13) finally results in

< 1 • (3.23)

For an evaluation of the reaction rate constant ~ and the di­

mensionless contact time e from the experimental results, the

gas-liquid interfacial area a and the contact time T must be

known.

40

0.20

0.15

l/H (kg/m3 ata)

100

75

(%)

50

25

0

Figure 3.15.

0.10

0.05

0 5.0 150 250

temperature (0 c}

Figure 3.14.

SoZubiZity of oxygen in water.

2---3----1

e

Effect of e on deviation of average absorption rate of oxygen

q from steady state absorption rate q • 0 . . o,s

41

The gas-liquid interfacial area in the continuous reactor can

be calculated from the gas fraction E (0.20 at 190°c), evalu­

ated from the residence time distributions, and the average

bubble diameter db (16 mm at 190°c). The bubble diameter was

estimated from bubble frequencies, evaluated from conductivity

measurements inside the bubble column.

The specific surface area in the continuous flow reactor

defined by

6£ ac = db

a , c

equals 75- 1 • From the physical absorption measurements dis­

cussed in 3.3.1. it followed that

-4 So the overall mass transfer coefficient K1 = 5.5 x 10 m/sec.

This is also the value of the liquid-side mass transfer coef­

ficient k 1 , if the gas-phase resistance is neglected.

According to Higbie's theory (118), k 1 is related to T by

k = 2 ,G_ 1 v n:r· (3.24)

The diffusivity is calculated from the tabulated dependence of

the viscosityµ on temperature (128), assuming

Dµ = T

constant.

-8 2 With the value of D0

of 2.3 x 10 m /sec and the calculated

value of k1

it follows from equation (3.24) that T ~ 0.1 sec.

Now we have calculated a and T for the continuous flow exper­

iments. For the semi-batch experiments we have to do the same,

but the values will be only estimations since the physical ab­

sorption rate was not measured in this system.

42

In the semi-batch experiments which were used for the evalua­

tion of k oxygen was passed over the solution. The available

area would be about two to three times the cross-section of

the reactor (A), because the surface would be somewhat dis­

turbed by the shaking action. Since the liquid hold-up (V) was

0.650 x 10-3 m3 the specific surface area per unit liquid vol­

ume (ab} is assumed to be

a -b -2.5 A = 20 m-1

v

The shaker moved every 0.4 sec. Therefore it was assumed that

the contact time T was of the order of 0.4 sec.

The actual value of T will also depend on the physical con­

stants of the solution, like viscosity, surface tension, and

density.

Because the order of magnitude of the gas-liquid interfacial

area and of the contact time are determined, the reaction rate

constant k can be estimated.

After calculating ~' the condition

(3.23)

must be checked. If this condition is fulfilled, equation

(3.19) may be applied.

In the computations it was assumed that the Henry coefficient

and the diffusivity of oxygen in water were independent of the

dissolved organic material. Figure 3.14 shows the reciprocal

Henry coefficient and its dependence on the temperature. The

data were taken from Pray (78) and Battino (113).

Table 3.2 shows the results of evaluations from semi-batch ex­o

periments with non-preheated glucose at 178 and 200 C and from ·o continuous flow experiments at 178, 190 and 200 C. It follows

from this table that condition (3.23) is only fulfilled in the

43

Table J.2.

Estimation of k.

0 temp. ( C)

[ 0. ] 1

(kg/m3 )

[cb] (kg/m3 )

ls -1 (sec )

a ( - )

~T( - )

2 [oi]

[cb]

continuous flow

178 190

0.31 a.so

24 15

0.1 0.8

0.8 2.4

0.01 0.08

0.026 0.067

exp. batch exp.

200 178 200

o.so 1. 22 1.43

10 12 12

2.1 0.3 1.0

4.0 1.2 3.2

0.2 0.12 0.4

0.1 0.2 0.24

continuous flow experiments at 190 and 200°c and in the semi­

batch experiments at 200°c. So, equation (3.19) may be applied

only to these experiments.

From the above it may be concluded that the calculation of the

reaction rate constants from the continuous flow experiments

at 190 and 200°c and from the semi-batch experiment at 200°c

can be considered a correct procedure. So, the reaction rate

constant is about 2 sec- 1 at 200°c and 0.8 at 190°c.

For the other experiments it is, therefore, necessary to know

the influence of e on the absorption rate. This influence was

already shown in figure 3.15, where the per cent. deviation of

the average absorption rate from the steady state absorption

rate for e + oo is given as a function of e.

44

It follows that the absorption rate decreases appreciably with

e as long as e < 2. Since 8 contains [cb]/[oi], e will also

depend on M which is

Therefore, in the case of continuous flow experiments at low

temperatures, where e < 1, it is to be expected that with de­

creasing values of M the conversion increases more than is

predicted with the model for the macro kinetics, based upon

equation (3.19}.

As may be seen from figure 3.12 the experiments at 175 and

186°c seem to confirm this conclusion.

45

f f I

• ' I I

I I I I r I • I 1 I 1 • I I I

I I I

1 I I - ..... ., "",_.. .............

Chapter 4 THE DISSOLUTION OF SLUDGE

In chapter 2 it was stated that sludge particles dissolve at

elevated temperatures by hydrolysis. Consequently, wet-air ox-:

idation may proceed through direct oxidation of sludge parti­

cles and through oxidation of hydrolysis products. In order to

understand the contribution of oxidation of hydrolysis prod­

ucts to the total oxidation rate, the degree to which the

sludge can be hydrolised into soluble products and the rate of

hydrolysis were studied.

4.1. APPARATUS AND EXPERIMENTAL DETAILS

Experiments were carried out in two batch-wise operated cylin-·

drical stainless steel reactors with volumes of 22 ml (~23x25 ·•

mm) and 52 ml (~34x38 mm).Each reactor was closed with a swiv-:

el. A thermocouple was attached to this swivel, and by closing,

46

the reactor, the thermocouple was placed inside. The normal

procedure was to fill the reactor with 15 ml of sludge and

then to heat the reactor in a glycerine bath of 60 - 90°c.

above the desired temperature, which latter was reached in

about 20 seconds. In order to secure a good heat transfer from

the oil bath into the reactor, and to keep the sludge parti-.

cles in suspension, the reactor was shaken with a flask-shaker

in vertical direction at a frequency of about 1000 min- 1 • When

the desired temperature was reached, the reactor was moved

over to a thermostat filled with Nassa oil. After the desired

residence time in the thermostat the reactor was cooled in a

water bath to 75°c, which took 15 seconds. The reactor was

opened and the contents were filtered on a heated Buchner fun­

nel.

Experiments were carried out with primary and activated sludge,

which had been obtained from the sewage works at Eindhoven.

The raw primary sludge contained much fibrous matter; fibres

with a length of 3 to 5 cm were observed. The fibrous nature

of this sludge made it impossible to take reproducible.charges

for the reactor. In order to overcome this difficulty, the 0 sludge was frozen with liquid nitrogen (-196 C) and was ground

in a marl mill. By means of a microscope it was found that the

grinding resulted in pieces with a length of 0.1 to 1 mm. The

sludges were stored at -20°c. No effect of freezing, grinding

and storage on the outcome of the dissolution experiments was

observed.

4. 2. THE HYDROLYSIS OF SLUDGE PARTICLES

Figure 4.1 shows characteristic examples of the course of the

hydrolysis process of activated sludge at 230°c. The COD of

the hydrolised matter,[c], is indicated as a function of time.

It follows from this figure that the curves can be divided in­

to two stages:

47

15

[C:J

10

(kg/m3 ) CJ

pH COD

+ 2.5 52 5 7.4 59 0

l!I 7.2 28

2 4 6 8

time (min)

Figure 4.1.

HydroZysis of activated sludge at 2J0°c.

(i) The first stage represents the actual hydrolysis process.

48

In general, a hydrolysis is a reversible reaction and an

equilibrium is reached between solids and hydrolised prod­

ucts when carried out batchwise. This seems also to occur

in our experiments.

From the repeated- hydrolysis experiments which will be

discussed in 4.2.3., it follows that several groups of

solids can be distinguished in the sludge which differ in

hydrolysis properties. The largest contribution to the

concentration of hydrolised matter results from a group

of solids which is nearly completely hydrolised in the

first hydrolysis step (see 4.1.3), while the other groups

only give a small contribution. Therefore, the hydrolysis

rate is largely determined by the first group and may be

described as an irreversible reaction.

This allows the rate of hydrolysis to be given as

d rAAl d[C:s] ~ k_[cs] dt ~ - dt = (4.1)

in which [e] is the COD of the hydro,lised matter, [csJ the

COD of the first mentioned group of solids, and ~ the

rate constant of the hydrolysis process. Since

[e J ~ [c J - [cJ s max ( 4. 2)

in which [c ] is the maximum value of [c], equation max (4.1) can be transformed into

~ = k([e J -[cJ). dt - max (4.3)

The results presented in figure 4.1 are plotted in figure

4.2 in the way indicated by the solution of equation

(4.3).

1.0

0.6

0.4 [cmax]-[c]

[cmaxl 0.2

0.1

0.04

Figure 4.2.

+, pH 'G.o +2.5

8

~o7.S

8

+ 0 7 .2 0

°\0

2 +

~ (min- 1 ) 1.1

0.6

0.9

0

4 6

time (min)

First order hydrolysis rate of aativated sludge.

From this graph it follows that equation (4.3) describes

the hydrolysis rate and that~· is of the order of SO h-l.

Takeichiro Takamatsu (120) studied the hydrolysis of ac­

tivated sludge at about 200°c. Although he took the first

49

sample after 10 minutes, he could calculate from his data

that the hydrolysis rate constant had to be larger than

36 h- 1 . This is of the order of magnitude of our value.

(ii) In the second sta.ge the concentration of hydrolised mate­

rial is constant or decreases slightly.

Takeichiro Takamatsu (120) found a constant level of the

COD of the hydrolised matter, while Brooks (23) observed

a small decrease. Brooks demonstrated that the hydrolised

material was again partly converted to the solid state.

The micro-organisms in activated sludge have dimensions of the

order of 1 - S µ. Because of their slimy skins the organisms

agglomerate into loose and porous structures with dimensions

of SO to 100 µ, known as flocks.

The rate constant associated wi.th the transfer of hydrolised

sludge from the outer surface of the flocks into the bulk of

the continuous water phase equals kfaf, in which kf is the

mass transfer coefficient on the outside of the flock and af

is the specific outer surface.

If we assume uniform spherical flocks with diameter df' then

the value of kf for these small flocks is given by

Sh

where Dr is the diffusivity of the hydrolised material and Sh

the Sherwood number. With the following characteristic values:

df = 60 x 10-6 m (87) and Dr= 10-8 m2/sec (diffusivity of sug­

ar molecules in water under experimental conditions), it fol­

lows that

-4 kf ~ 3.3 x 10 m/sec.

The suspension contains 4 volume per cent. of solids and the

flocks consist for at least 60% of water (87). Therefore, the

so

volume fraction of flocks (ef) is larger than 0.10. The spe­

cific outer surface is given by

Thus, -1 -1 kfaf ~ 3.3 sec = 11,900 h •

Owing to the small dimension and the loose and porous struc­

ture of the flock, the rate constant of physical transport in­

side, will also be of the order of 11,900 h- 1 •

Since the experimentally determined rate constant of the hy­

drolysis process is of the order of 50 h- 1 , the process must

be limited by chemical reaction inside or on the surface of

the solids.

The reaction~limited hydrolysis implies that the concentra­

tion of hydrolised matter inside the flock is uniform and

equals the concentration in the bulk of the liquid in which

the flocks are suspended,[c]. When the reaction is interpreted

as a homogeneous reaction this further implies that the rate

of hydrolysis can be described by

In this, ~h is the hydrolysis rate constant and Es is the vol­

ume fraction of solids in the flock.

The experiments showed that

Using the characteristic values for Ef ~ 0.1 and Es ~ 0.4 it -1 follows that ~h ~ 12,000 h •

The hydrolysis rate of primary sludge was determined at 200°c

51

0.8

0.6 200°c

0.4 0 pH = 1.0

COD0

= 37 kg/m3

le J-lcl max

0.2

l 2 3 4 0.1 ""-~~~-----'--~~~--&.~~~~~"--~~~~--~~------'

time(min)

Figure 4.3.

FiPBt OPder hydrolysis Pate of primary sludge.

and at pH = 1.0. The results are given in figure 4.3 where

[cmax] -[c] log [c ]

max

is plotted as a function of time. From this graph it can be

concluded that the hydrolysis of primary sludge is also a

first order process~The calculated rate constant of the hydro­

lysis process has a value of 30 h- 1 .

Since the ground primary sludge particles have characteristic

dimensions which are only somewhat larger than those of the

activated sludge flocks, the mass transfer from the outer sur­

face of the particles will not be of influence on the outcome

of the experiments either.

52

This section deals with the hydrolysis of the residue ob­

tained after the first hydrolysis step.

Repeated hydrolysis was studied with primary sludge at 200°c

as a function of the initial concentration of the suspension.

Results of the first hydrolysis step are indicated in figure

4.4 by curve (1) where the maximum concentration of hydrolised

material [emax] is plotted as a function of the initial total

COD of the sludge in the first step. The residue was filtered

off and was again suspended in water to the original volume

and [e ] was determined.These results are indicated by curve max {2) and are also plotted as a function of the total COD of the

sludge in the first step. From the residue obtained after the

second step [emax] was also determined; it is indicated by

curve ( 3).

It follows from figure 4.4 that the first hydrolysis step

provides the highest value of [emax]' and that in the subse­

quent steps lower values are obtained which are nearly inde­

pendent of the initial sludge concentration. Apparently the

10

200°C

8

0

0/ pa= 1.5 /

7 min ~ 0

(1)

[cmaxj

/0 4

+ ----------:-+ - ( 2)

2 +

-t:l t:l-1::]- ( 3)

0 10 20 30 40

COD of starting sludge (kg/m3 ) Figure 4.4.

Repeated-hydrolysis of primary sludge.

53

sludge cannot be considered as one hydrolisable material.

The second and third steps indicate that the sludge contains a

group of hydrolisable solids which results in sparingly sol-

uble products,by which the concentration*) becomes independent

of the initial sludge concentration as long as this group con­tains enough material to build up that concentration.

The first step shows that the sludge also contains a group of components which are almost completely hydrolised in this step.

From the hydrolysis rate experiments it followed that the rate

constant of this group is about 30 h-l at 200°c (see 4.2.2.).

It is possible that completely unhydrolisable material is also

contained in the sludge.

4.3. INFLUENCE OF CONCENTRATION AND OF OPERATING CONDITIONS

ON [cmax]

The influence of concentration and of operating conditions was

investigated in the first hydrolysis step. As followed from

4.2.3., in this step [cmax] is mainly determined by a group of solids which almost completely hydrolise in this step.

(i) Influence of sludge concentration

The influence of the sludge concentration followed al­

ready from the repeated- hydrolysis experiments for pri­

mary sludge at 200°c. The influence under other opera­

ting conditions for both primary and activated sludge

are shown in figure 4.5. This graph shows that [cmax] increases more or less linearly for both types of sludge

and that [cmax] is about the same, viz. 25 per cent. of the total concentration of the suspension.

(ii) Influence of temperature

54

Table 4. l shows the i.nfluence of temperature on [ cmax]

for activated and primary sludge. It is seen from the

table that in the investigated range of temperatures

[cmaxl only slightly increases with temperature.

*) Concentration of dissolved matter

15

[ cmax ]

(kg/m3) 10

5 ...

0

Figur>e 4. 5.

pH

+prim.sl., 4

oact.sl. 7

25 so 7,5

temp

(OC)

230

230

COD of sludge (kg/m3 )

Effect of initial sludge concentr>ation on [a ]. max

Tab le 4. 1.

Influence of temper>atur>e on [a ]. max

activated sludge primary sludge

[co] = 65 kg/m\ pH = 6.8 [co] = 30 3 kg/m ; pH= 1.5

Temp. [cmax] Temp. [cmax]

( oc ) ( kg/m3 ) ( oc ) 3 ( kg/m )

205 17.0 200 8.0

225 17.7 290 11. 5

250 18.4

(iii) Influence of pH

During wet-air oxidation the pH decreases from 7 to a

value between 6 and 4 as a result of the formation of

acids.Therefore the influence of pH on [c ] was inves-max

55

56

tigated for activated sludge at 230°c, while the COD of

the sludge was 65 kg/m3 •

The pH was set with sulfuric acid. Table 4.2 shows the

results. It can be concluded from the data that the pH

has a slight influence on [cmax] in the range of experi­

mental conditions.

Tab'le 4.2.

Inf'luenae of pH on [8 J of activated s'ludge at 230°~ ma.x

pH 4.5 5.3 6.1 6.9

[cmaxl (kg/m3) 16.6 16.2 15.6 15.2

' •

Chapter 6 THE OXIDATION OF PRIMARY SLUDGE

5.1. PRELIMINARY BATCH EXPERIMENTS

Preliminary batch experiments were performed with primary

sludge in the 52 ml reactor, described in 4.1. The reactor was

charged with 15 ml of ground sludge and pure oxygen.

The pH was set with sulfuric acid at 4. The repoJ."'lted oxygen

pressures were measured at room temperature before the exper-3 iment. At a partial pressure of 25 atm, 90 kg of oxygen per m

of sludge suspension was present.

The charged reactor was preheated, kept at 230°c in the ther­

mostat, and cooled, as described in 4.1.

Figure 5. L shows· the influence of the oxygen partial pressure

on the conversion for several sludge concentrations after a

reaction time of 3 min. It follows from this graph that there

is a "critical" pressure~

57

40

30 COD

reduction (%)

20

10

Figure 6.1.

--++----

---~~0--~~0

r ['J

['J COD 230°c 0' 0 I

0 4. O kg/m3 ['J ['J pH = 4

j 0 28 .1 II

+ 32. 6 II 3 min .

10 20 30

oxygen pressure (a ta)

Effect of oxygen pressure and siudge concentration on

oxidation of primary siudge.

Below this pressure the relative COD reduction increases line­

arly with oxygen pressure (conversion order in oxygen equals

one); above it, the relative COD reduction is independent of

the partial pressure (conversion order in oxygen is zero) .

Furthermore it follows from the graph that below the "criti­

cal" oxygen pressure the COD reduction is independent of the

sludge concentration which suggests that the conversion order

in COD is one, while above it, the relative COD reduction in­

creases with COD.

In practice the oxygen pressure is of the order of 5 to 10 atm

which means 2.5 to 5 atm when measured at room temperature,

which is far below the critical oxygen pressure.

Further experiments were carried out in which the residence

time of the reactor in the thermostat was varied. The sludge

had a starting COD of 24.4 kg/m3 while the pH was set at 4.0.

The oxygen pressure, measured at room temperature before the

58

1. 0 ~ ,, ,, ,,,

pH = 4. 0 \ \ \ \ 230°c \ \

COD \ \ COD0

= 24.4 kg/m3

\ \ COD0 \ \

' ',(1)

0.5 '(2)

( l) Po = 5 atm 2

( 2) " = 10 II

( 3) II = 29 II

0 2 4 0 time (min)

Figure 5. 2.

Effect of time on oxidation of primal>y sZudge.

experiment, was fixed at 29 atm, which is above the "critical"

value. The results are shown in figure 5.2. It is seen from

this graph that it took about a minute before noticeable ox­

idation occurred.

The same phenomenon was found in the oxidation of the model

sludge (see 3.1.2.). The experiments discussed in 5.2.3. dem­

onstrated that by preheating the sludge for five minutes, the

"lag phase" disappeared. This indicates that some "thermal"

reaction had to take place for the oxidation of primary sludge,

too.

5.2. SEMI-BATCH EXPERIMENTS

From chapter 4, where the hydrolysis of sludge particles was

discussed, it followed that at least 30% of sludge passes in-

59

to solution. This section deals with the conversion rate of

hydrolised sludge and of sludge as such, where oxidation of

hydrolised sludge and sludge particles takes place simulta­

neously.

The oxidations were carried out in the one-litre autoclave de­

scribed in 3.1.1. In this reactor 250 ml of water was heated

to the desired temperature. Then 250 ml of ground primary

sludge was injected with nitrogen. After 5 minutes the desired

temperature was reached again. The nitrogen present was quick­

ly replaced by air and the air flow was set.

The solutions of hydrolised sludge were prepared in a half­

li tre autoclave, also provided with a shaker (see fig. 5.3).

In this autoclave, 100 ml of water was heated. Then 250 ml of

sludge was injected with nitrogen. After three minutes at

60

air

injector

filter hydroliser

Figure 5.3.

cooler

oxidiser

spent gas

cooler

samples

Apparatus for o~idation of hydroZised sludge.

230°c, the solution thus obtained was pressed through a stain­

less steel filter placed inside the autoclave. The freshly

prepared solution was driven into the one-litre autoclave, in

which 150 ml of water had already been heated to the desired

temperature. Five minutes after injection of the solution of

hydrolised sludge the nitrogen was replaced by air and the air

flow set. The oxidations were performed at 230°c and 100 atm.

The air flow was fixed at 1 Nm3/h. The frequency of the shaker

in the oxidation autoclave was 144 min- 1 • Owing to the exper­

imental conditions, the shaker forced gas bubbles from the gas

phase above the suspension into the sludge.

The scheme of experiments carried out is shown in figure 5.4.

(i) Sludge was oxidised as such. Since the sludge was pre­

heated for five minutes, hyGrolysis was started already

before air was fed into the system. This means that from

the moment that air was introduced into the reactor, simultaneous oxidation of hydrolised sludge and sludge

particles occurred (overall reaction rate constant ~t).

(ii) . The sludge was also hydrolised and filtered.

The solution obtained was oxidised (reaction rate con­

stant ~d).

(iii) The residue of the filtration was again suspended in

water and oxidised (reaction rate constant k ). Since . -r this suspension was also preheated for five min, simul-

taneous oxidation of hydrolised residue and residue par­ticles occurred from the beginning.

(iv) The residue was also hydrolised and filtered,while next,

the filtrate obtained was oxidised (reaction rate con­

stant !sf) •

In order to obtain relatively high concentrations of hydro­

lised sludge and hydrolised residue, the sludge used to pre­

pare the solutions for the experiments (ii) and (iv) had a

higher concentration than the sludge used for experiment (i) •

61

sludge sludge oxidise

~t •product I

dissolve

,vdrolised oxidise filtre ! sludge ~d

.product II

residue

!suspend suspended oxidise lin water residue ~r

product III

suspended residue

dissolve

ovdrolise1 oxidise

I .. filtre residue !sf product IV

·- product v

Figure 5.4.

Soheme of experiments and produote.

All experiments were performed at pH= 4. Furthermore, the ex­

perim~nts (i) and (ii) were also studied at pH's 1.5 and 7.

During experiments the pH changed, depending on the initial

value. pH 1.5 changed into 2.0, 4 into 3.3 and 7 into 5.1.

From the preliminary experiments discussed in 5.1.,it followed

that the COD reduction after a reaction time of 3 min is first

62

order in sludge concentration and first order in oxygen pres­

sure at the oxygen pressures applied in practice and also in

our further investigation.

Since the maximum COD reduction generally is less than 100%,

the conversion rate cannot be fully described by the assump­

tion of a first order in the total COD. By the oxidation of

glucose (Chapter 3) the effective COD was introduced, being

the actual COD minus the COD at the point where the conver­

sion rate is nearly zero.

The starting point of the model discussed now is the assump­

tion that the chemical kinetics of the oxidation are described

by the following rate equation (g).

in which k is the second order reaction rate constant, [c] the

effective COD and [ob] the oxygen concentration in the bulk.

The conversion rate is partly limited by the transport of ox­

ygen from the gas phase into the suspension. It is assumed

that the gas phase and the suspension are completely mixed.

It should be remembered that the gas bubbles in the suspension

originate from the gas phase above it.

When neglecting the gas phase resistance for mass transfer,

the oxygen pressure at the gas-liquid interface (pi) becomes

identical with that in the spent gas (pe).

The following mass balances were obtained, in which ¢0

x is the

amount of oxygen passing through the gas-liquid interface per

unit time and per unit reactor volume, and which is given by

¢ = k a ( [ o1. ] - [ob]) ( 1-e ) ox t

where [o.] is the oxygen concentration at the interface which 1

is related to p. with the Henry coefficient H by 1 .

= p /H. e

63

E is the gas fraction, based on the total reactor volume.

The mass balance over the sludge phase leads to

where

k[c] [o ] = - £1£1 - b dt

The mass balance over the gas phase is

This can be written as

in which G is the total volumetric gas flow under reaction

conditions, v is the volume of the reactor and pf is the ox­

ygen partial pressure in the gas feed, after saturation with

steam.

Introducing the dimensionless concentrations and pressure

Y = lsL [ C ] I

0 I p =

where [c0

] is [c] at zero time, and eliminating

mass balances, the following equations are found

- .£y = K l"ly dt 2,...

~ = K~ [ K2py - Kl (n-p) ]

d1T 1 - dt = Ks [ Kl (n-p) - K4 (1-n) ]

64

¢ from the ox

In these equations

Kl kta pf -1 = [time ] [c ] H

0

K2 !s pf -1 = -H- [time ]

K3 Pf

= [c0

]H

K4 G pf -1 = [time ] VRT [c ) (1-e:)

0

KS pfE

= [co] (1-e:) RT

The boundary conditions are obtained from the values of y,p

and TI at zero time Then y = 1, and p = 0. The value of TI de­

pends on the experimental procedure.

Near the gas-liquid interface there is a transfer film or

diffusion film where the concentration of oxygen is higher

than the concentration [ob] in the bulk of the liquid.

The oxidation rate in this film, will therefore also be higher

than in the bulk. When,however, the thickness of the film is

small enough, the contribution of the film to the overall ox­

idation can be neglected. The condition on which this is per­

missible runs

k[C)/D = - 0

2 k [CJ /kt < 1. t/J = 6 v Here 6 is the thickness of the diffusion film related to kt

by kt = D0/6. t/J 2 equals the ratio of the maximum oxidation

rate in the film(= k [o.][c]6) and the maximum diffusion rate - l.

through the film (= D [o.]/6). 0 1

This condition has to be checked after evaluation of k.

65

Figure 5.5 gives a typical example of the oxidation of sludge.

The CODs of the suspension and of the dissolved matter are

given as a function of time. As a result of the hydrolysis

during preheating the COD of dissolved matter is relatively

high at zero time. It can be seen from the graph that this

concentration increases rapidly during the first few minutes

of the oxidation. This might be caused by soluble oxidation

products resulting from partial oxidation of the solids. It

also follows from the graph that the "lag phase" has disap­

peared through preheating.

10

COD

(kg/m3 )

5

0

Figure S.S.

+\ +

230°c

100 atm

pH = 4

\sludge {product I)

'+. r~~+ filtrate ~-~---

5 10 15

time (min)

Typical example of semi-batch oxidation of primary sludge.

The results of oxidation of sludge, solutions, residue and so­

lution of residue are represented in table 5.1 and are partly

shown in figures 5.6 and 5.7.

In these graphs the dimensionless effective COD is plotted

against time.

66

The results were analysed with the model presented in 5.2.2.

The boundary conditions associated with the experimental pro­cedure are

t = 0 y = 1

7f = 1

p = 0

The constants K3 , K4 and K5 are known. The,order of magnjtude

of K1 , the transfer term, can be estimated.

Table 5.1.

COD in kg/m 3 of pPodua~s afteP tPeatment as indiaated

in figuPe 5.4.

Reaction time in min

Product p.H 0 1 2 3 4 5

total 7.0 13.3 11.9 10.0 9.3 - 7.7

sludge 4.0 10.4 8.2 6.5 6.0 - 4.9

(product I) 1.5 14.5 11. 4 9.3 7.2 - 5.3

hydrolised 7.0 6.2 5.4 4.1 4.1 3.8 3.3

sludge 4.0 7.0 5.4 5.0 3.9 3.6 3.1

(product II) 1.5 7.3 5.6 4.5 3.6 3.0 2.3

suspended 4.0 5.8 5.3 4.8 4.5 4.2 -residue

(product III)

hydrolised 4.0 3.0 2.7 2.5 2.0 1. 7 1.6

residue

(product IV)

From Higbie's theory it follows that

suming a contact time of 0.1 sec and 3.3.4.). In 3.1.3. it was pointed out

7 9 12 15 20

6.2 5.9 4.9 4.0 3.4

4.1 - 3.2 2.6 1.9

4.0 3.8 3.1 2.8 2.6

2.8 2.4 2.4 2. 3 . -2.5 1. 9 1.9 1. 7 -1.6 1. 3 1.2 1.1 -- - 3.1 - 2.3

1.2 1.0 0.9 0.8 -

k R, 6 -4

c:: x 10 m/sec, -8 2

Do = 3 x 10 m /sec that a c:: 270 m- 1 •

as-

(see

The reaction term 'K2 was evaluated from the experimental re­sults with the aid of an analogue computer.In the calculations

it was assumed that maximum conversion was reached when the

last sample was taken. It can be seen from table 5.1 that this

assumption will not always be fulfilled,which might cause some

67

trouble in the description of the conversion near the end of

the oxidation.

The calculated values of k and w are shown in table 5.2. It

follows from the table that w << 1, which means that the re­

action within the diffusion layer around the gas bubbles can

indeed be neglected.

Table 5.2.

Estimated values of k and ¢ • - max

Product pH ~(m3 /kg h) ¢max

total 7.0 17 0.052

sludge 4.0 21 0.051

(!st) 1.5 28 0.088

hydrolised 7.0 22 0.040

sludge 4.0 23 0.042

(~d) 1.5 31 0.053

suspended

residue 4.0 12 0.027

(~r)

hydrolised

residue 4.0 17 0.023

(}Sf)

The measurements and calculated curves are presented in the

figures 5.6 and 5.7.

Figure 5.6 shows that the conversion of hydrolised sludge is

well described by the model. Consequently, the values of ~d

will be reasonably accurate.

68

1.00 0 pH = 7

0

0 pH = 4 0.75 + pII = 1. 5

y

' 0.50 +\ 230°c

Oo 100 atm ·\ \0

0.25 \\ +\""'

0 5 ~~0-time (min)

Figure 5. 6.

Oxidation of hydrolised sludge (product II).

Data points and computerised curves.

0.70

y

0.50

0.25

0

Figure 5.?.

230°c

100 atm

0 pH = 7

O pH = 4

+ pH = 1. 5

time (min)

20

Oxidation of primary sludge (product I).

Data points and computerised curves.

69

From figure 5.7 it can be concluded that the model is too

simple for the oxidation of sludge as such since it hardly de­

scribes the course of conversion as function of time.

In the next chapter, where the oxidation of activated sludge

is discussed, the model is modified thus that a simultaneous

oxidation of two groups of components in the sludge which

differ in reactivity is assumed.

However, the values of ~t given in table 5.2 are first esti­

mations which will be of the order of magnitude of the real values.

From the values of ~ reported in table 5.2 the following con­clusions can be drawn.

(i) The overall reaction rate constant, ~t' based on the ef-'

fective COD of the suspension, [ct], is almost constant

in the pH range encountered in wet-air oxidation

( 4 ~ pH {- 7).

(ii) The reaction rate constant ~d of hydrolised sludge

(product II) is almost independent of the. pH in the

range 1.5 ~ pH ' 7.

(iii) In the range 1.5 E pH~ 7, ~dis cf the order of ~t·

(iv) The reaction rate constant of suspended residue (product

III) is lower than that of the sludge. This could mean;

that the sludge consists of different groups of compo­

nents which vary in reactivity. Apparently the more re­

active material has largely disappeared in the foregoing

dissolution step.

(v) The reaction rate constant of hydrolised residue (prod­

uct IV) at pH= 4 is somewhat smaller than that of hydro­lised sludge (product II).

70

Even though the residue (product III) is less reactive,

it contains some reactive hydrolisable matter.

Easily soluble matter is apparently also easily oxidised.

The total conversion rate (gt) is composed of contributions of

dissolved matter <gd) and of solid matter (gs).

It follows from the experiments that the reaction takes place

in the bulk of the suspension. So,

Furthermore,

Analogously it can be assumed that

R = k [cs] [ob] , -s -s

in which ~s is the reaction rate constant of solid matter and [cs] the effective COD of solid matter.

By introducing the fraction ~ of dissolved effective COD, de­

fined by

[cd] ~ = TCT I

t

the equations can be transformed into

The value of ~ depends on hydrolysis and oxidation.

From figure 5.5, which illustrates a typical example of the

paths of [at) and [ad] during an experiment at pH = 4, it fol­

lows that ~ first increases rapidly and then remains practi­

cally constant in the course of the experiment.

As a first approximation ~ may be considered constant at a val­

ue of o.6.

71

.70 0 0

JJ--· 0-~· 0 .

0 ., .so

.30

·10

0 20 40 60

initial COD {kg/m3 )

Figure 5. 8.

Fraction of dissolved COD after 4 min as a function of

initial COD.

The value of ~ also follows from batch experiments which were

not discussed in section 5.1.

These experiments were carried out at 230°c, pH = 4 and an

oxygen pressure of 25 atm. After a reaction time of 4· min the

CODs of the resulting sludge and of the filtrate were deter-

mined. This was repeated for several initial CODs The re-

sults are shown in figure 5.8, where the ratio of the COD of

the filtrate {[cd]) to the COD of the resulting sludge {[ct])

are plotted against the initial COD. It can be seen from this

graph that the ratio is almost independent of the initial

sludge concentration at a value of about 0.6. Since

it can be concluded from these experiments too that ~ ~ 0.6.

Substitution of ~' ~t and ~d in equation {5.1) yields for

pH = 4.

72

k = 18 m3/kg h. -s

This shows that the hydrolised sludge is somewhat more reac­

tive than solid sludge, but is not strongly deviating from it.

It can be concluded that to describe the total conversion rate

at least at pH = 4, and probably also at the other pH's it is

not necessary to take the hydrolysis into account as a sepa­

rate reaction step.

73

Chapter 6 THE OXIDATION OF ACTIVATED SLUDGE

The experimental results of the oxidation of activated sludge (see e.g. figures 6.2 and 6.3) were analYsed with the con­version model presented in chapter 5. It was found that it was

not possible to describe the complete curves with one reaction rate constant. The oxidation of primary sludge, discussed in

chapter 5, suggested already that more than one reaction rate

constant is necessary for a fair description of the conversion. Therefore,the model was extended to the simultaneous oxidation of two groups of components which differ in reactivity.

6.1. EXTENSION OF THE MODEL

The total concentration [e] of the sludge is supposed to be

the sum of the concentrations [a], [b] and [c00], which are the

concentrations of a more reactive group, of a less reactive group, and of a non-reactive group,respectively.

74

Hence,

[c] = [a] + [b] + [c ] • 00

By introducing the dimensionless COD of the sludge (y) which

is the actual COD ([c]) divided by the initial COD ([c ]) , it 0

follows that

( 6 • 1)

in which a = [a]/[c0

] and B = [b]/[c0]. The maximum COD reduc­

tion that can be obtained (f ) is max •

Thus

-y = a + B + 1 - f max ( 6. 2)

-At zero time y = 1. Substitution of this in equation (6.2)

results in

a + B = f o o max

where a 0

is the fraction of more .reactive material

sludge and B0

is the fraction of less reactive material.

( 6 • 3)

in the

The equations describing the non-stationary conversion rate in

the semi-batch system are very much analogous to those derived

for the oxidation of one group of components (see 5.2.2.).

Only the mass balance across the sludge phase must be modified

to

75

while

- ~ = k [a] [o ] dt -a b and

in which ~a and ~b are the reaction rate constants of the more reactive group and the less reactive group respectively.

The other mass balances remain unchanged.

After introduction of the dimensionless concentrations in the

mass balances, the following set of simultaneous differential

equations is obtained.

da. K a.p = dt (.l

- df3 = Kf3f3p dt

2:12. l (K(.la.p + Kf3f3p - Kl (TI-p)] = dt K3

dTI l (Kl (TI-p) -K(l-TI)] = dt KS 4

in which

TI,p, K1 , K3 , K4 and K5 were introduced in 5.2.2.

Since the experiments were started with a nitrogen

above the sludge, while at zero time air was added

driving out the nitrogen, the initial conditions are

t = 0 a. = a. 0

f3 = f max - a.o TI = 0

p = 0

6.2. APPARATUS AND EXPERIMENTAL DETAILS

atmosphere

gradually

Experiments were carried out in the one-litre autoclave de~

76

scribed in 3.1.1. In this reactor 150 ml of water was heated

to the desired temperature and 350 ml of sludge was injected

with compressed nitrogen. After 3 - 4 minutes the desired tern-

perature was reached again

determine the initial COD.

was added, which gradually

and the first sample was taken to

Five minutes after injection, air

drove out the nitrogen. The oxygen concentration of the spent gases was continuously measured and·

recorded. As a function of time, samples were taken from the sludge phase, in which the COD was determined.

Experiments were carried out at temperatures ranging from

180° - 290°c and pressures ranging from 43 - 150 atm. The air

flow was always 0 . 3 7 Nm 3 /h. ·The shaker frequency was 14 4 min -1,

so that the gas-liquid interface was about 270 m2/m3 of sus­pension {see 3.2.2.).The experiments were started at the orig­inal pH of the sludge which ranged from 7 to 8.

Owing to the preheating the pH decreased to 6, while during

oxidation the pH further decreased to a value between 5 and 6.

6.3. EXPERIMENTS WITH SLUDGES OF DIFFERENT ORIGINS

Experiments were conducted with activated sludges produced at seven sewage treatment plants, viz., those of Eindhoven,

Amsterdam-West, Haarlem, Tilburg, Schijndel, Beverwijk and

Apeldoorn.These sewage plants treated waste waters originating

from highly industrialised cities to small villages which produce almost purely domestic waste water. The purpose of

these experiments was to examine the oxidation properties of

sludges from different origins.

By dilution

same COD of

or concentration the sludges were brought to the

20 kg/m3 and were oxidised at 240°c and 70 atm. owing to the experimental procedure the sludge was further

diluted inside the reactor resulting in an initial COD of 14

kg/m3 •

The results are shown in figure 6.1. From this graph it fol­lows that all sludges are oxidised in largely the same way.

77

d'11

240°c o Amsterdam-West

70 atm 0 Apeldoorn

COD

(kg/m3 )

4

Figure 6. 1.

0 ' I

• Schijndel

• Beverwijk

+ Tilburg

• Haarlem

x Eindhoven

* Eindhoven

40

time (min}

Semi-batch o~idation of activated sludges of different origin at 240°c and 70 atm.

This is a pleasant circumstance, since .it makes one reasonably

confident that the insight and conclusions to be arrived at

after further experiments carried out with activated Eindhoven

sludge, can be applied to any activated sludge at least in a

qualitative sense.

In the samples taken after 60 min

COD was determined. The results

the fraction of dissolved

are shown in table 6.1. It

follows from these data that after 60 min, when only 10% of

the original COD is left, the resulting COD consists for 92 -

98% of dissolved products. From the solid residue in these

samples the ash was determined by heating in air at 600°c to

constant weight (77).With some reservations the loss in weight

might be interpreted as the organic matter in the residue. The

78

results are also shown in table 6.1. From the data it follows

that the ash fraction is about the same for each residue with

a mean value of 10%.

This indicates that also in other respects these sludges are

comparable.

TabZe 6.1.

Properties of oxidised sZudge after 60 min.

COD sol /COD tot ash in solid

(%) residue (%)

Amsterdam 98 8

Apeldoorn 97 12

Schijndel 95 8

Beverwijk 95 14

Tilburg 96 11

Haarlem 92 9

Eindhoven 93 9

6.4. FURTHER EXPERIMENTS WITH EINDHOVEN SLUDGE

The effect of temperature was investigated in the range of 180 to 290°c. The corresponding pressures were fixed at those

values which provided at saturation the same oxygen concentra­

tion in the water phase of 0.58 kg/m3• This resulted in the following combinations of temperatures and pressures, given in

table 6.2.

initial COD was 13.4 kg/m3 . The

The COD

results are shown in figure 6.2, which gives the relative (y) as a function of temperature. It follows from this

graph that the temperature has a considerable effect on the

conversion rate.

79

Tab Ze 6. 2.

Temp (oC) 180 200 220 240 255 270

Steam pressure 10 15 23 32 42 55

(atm)

3 H (m ata) 0.042 0.050 0.061 0.072 0.088 0.111 kg

Oxygen pressure 13.7 11.5 9.5 8.0 6.5 5.2 (atm)

Total pressure 75 70 68 70 73 80 (atm)

•

1.00 computerised curves

o.75

y

0.50

0.25

0 40 • i 8

time (min)

Figure 6.2.

Effect of temperature on activated sZudge. Data points and

computerised curves.

80

290

74

0.131

4.4

95

It is remarkable that at 180°c after 90 min maximum conver­

sion is not reached, while according to the literature (70),

fma is about 50% at 180°c.In order to find f of our sludge x max at these relatively low temperatures,an experiment was carried out at 185°c and 75 atm during 7 h. The results are shown in

figure 6.3, from which is seen that after 7 h fmax is not yet

reached, but that fmax is of the order of 90% at iss0 c.

COD

(kg/m3 )

12

8

4

Figure 6. 3.

100

185°C

75 atm

-o'---o'---0-200 300 400

time (min)

Oxidation of aativated sZudge at 185°c and ?5 atm.

Since at 240°c f is 91%, it was assumed max and 220°c fmax is 90%. As follows from figure

fmax increases slowly with temperature to 93%

that at 180, 200

6.2, above 240°c

at 290°c.

6.4.2. Evaluation of kinetic data and discussion -----------------------------------------The experimental results were analysed with the conversion

model presented in 6 .1. .Since the constants K1 , K3 , K4 and K5 can be calculated, only Ka' K8 ,a

0 and 8

0 have to be evaluated

from the measurements. Since fmax follows from the figures 6.2 and 6.3 and since a

0 and 8

0 are related by

81

a + B = f o o max

either a0

or B0

is an unknown parameter.

The set of equations presented in 6.1.were simulated on an a­

nalogue computer. The procedure was to adjust the values of

a0

, Ka and KS in such a way that the conversion curve pro­

duced by the computer fitted the experimental curve as well as

possible. It will be clear that because of the spread and the

possible error in the experimental results, a , K and KB can o a also be determined only with a certain an)Ount of inaccuracy.

A mathematical error discussion of a0

,Ka and KB is rather com­

plicated because the four equations describe the course of

conversion simultaneously.For that reason preference was given

to an experimental determination of the error in a0

,Ka and KB.

Figure 6.4 gives for one single experiment the computer curves

belonging to different combinations (a0

, Ka, KS), which de­

scribe more or less the experimental results.

For 180, 240 and 290°c the ranges of values of Ka, KB and a0

which allow of a fair description, were determined. It was

found that, when a0

is fixed, only one combination of Ka and

KS is possible. The range over which a0

could be changed while

still a fair description was obtained, depended on the temper­

ature. The results are as follows.

At 180°c, 0.20 ~a ~ 0.65, at 240°c,

0 0.40 ~a ~ 0.70,

290°c 0

and at 0.65 ~a ~ 0.75. 0

For each proper solution the following relations were found

between a , K and K0 • o a µ

82

R a ~ constant a o

(6.3)

CXO K Ka ex

(-) (h-1) (h-1)

0.75 0.70 26 3,8 . 2 0.65 28 5.1 y

3 0.45 33 7,2 4 0,40 39 8.8

1 2

- . I I - . I I

3 4

0 0 10 20 30 40

time (min)

Figure 6.4.

Computer aurves for severai aombinations of a , K and K0 o a µ

for 240°C and 70 atm.

constant (6. 4)

The constants depend on temperature and are given in table 6.3.

The physical background of the first relation (6.3) can be

easily understood. In the first period of the oxidation the

conversion rate is mainly determined by the group of more re-d-

active components. So,the conversion rate - d~ is proportional

to K a , which means that a good description of the conversion a o

rate can be obtained by several combinations of K and a , as a o

long as the product of Ka and a0

is kept constant.

83

Table 6.3.

Constants in equations (6.3) and (6.4)

temp Ka Cl. 0 KS

(OC) (h -1) f max -a 0

(h-1)

180 0.10 0.078

240 17 17

290 180 41

..

The second relation (6.4) is a property of the system that can

not be derived analytically.

From the foregoing it can be concluded that there are a large

number of solutions which are capable of describing the exper­

imental results.

When the conversion model is correct,however, there can be on­

ly one correct solution at each temperature.

In order to find this solution, additional information is nec­

essary. Therefore it was assumed that a is a property of the 0

starting material, which implies that a must be the same at 0

each temperature. The only value for a0

which satisfies this

is a0

~ 0.65. The combinations of Ka and KS are then fixed at

each temperature. The results are given in table 6.4.

The figures 6.5 and 6.6 present an Arrhenius plot of ~a and ~b'

respectively. It follows from these graphs that Arrhenius' law

describes the effect of temperature rather well and that the

activation energy is about 23 kcal/mol for both reaction rate

constants. This seems to confirm that the assumption of a con­

stant value for a0

is reasonable.

The curves in figure 6.2, which show the influence of temper­

ature, were calculated with the values of ~a and ~b given in

table 6.4.

84

Table 6.4.

Influence of temperature on k • kb and f -a- - ma:r:

temp k ~b -a f max

(oC) (m3 /kg h) (m3 /kg h)

180 2.9 0.35 0.90

200 9.0 0.97 0.90

220 24 3.6 0.90

240 43 7.9 0.91

255 180 12 0.92

270 250 20 0.93

290 500 22 0.93

1000 100

\

\ \ 0

0\

0 100 10

\ ~a ~b

(rn3 /kg h) (rn

3 /kg h) 0

\ 10 0 1

\ \ 1 1.8 2.0 2.2 0.1 1 8 2 0 2.

1000/T (oK-1) 1000/T (oK-1)

Figure 6. 5. Figure 6. 6.

Effect of temperature on k . -a Effect of temperature on ~b·

85

1.00

·r+ 0 .75 \ t points of -

o\li 0 data y

+ data points of TI p

0 50 270°c

80 atm

0.25 \ 0

"' 0----0 0 0 0 0 4 6CI

time {I:>in}

Figure 6. ? •

y,TI and p for oxidation of aativated sludge at 2?0°c and

BO atm. Data points and aomputerised aurves.

As an illustration of the dependence of TI and p on time, the - 0 calculated curves of y, TI and p at 270 c are shown in figure

6.7. The graph also contains the measured dependences of y and

TI. TI was calculated from the measurements of the oxygen con­

centration in the spent gases. From figure 6.2 it followed

that the model describes well the dependence of y. Figure 6.6

shows that the dependence of TI is also described by the model.

The effect of the pressure was investigated in the range of

43 to 150 atm at a temperature of 240°c. This means that the

oxygen pressure varied from 2.3 to 23 atm. The initial COD of

the sludge was 13.6 kg/m3 .The results are given in figure 6.8,

where y is plotted against time. This graph shows that the

pressure,as expected, clearly affects the conversion rate.

86

-y

1. 00

0.7S

a.so

0.2S

0 10 20

Figure 6. 8.

Eff eat of pressure on

1.00

0.7S

-y

a.so

0.2S

0 10

Figure 6.9.

240°c

30 40 so time (min)

oxidation of aativated sludge

20

computerised curves

240°c

30 40

time (min)

so

60

at 240

60

Calaulated effeat of pressure on oxidation at 240 °c.

o-

0 c.

87

The effect of pressure was also calculated with the conversion

model. The values of lsa and lsb were taken from the experiments concerning the influence of temperature. The predicted effect

of pressure is shown in figure 6.9. A comparison between cal­culated and measured curves reveals that the predicted effect

of the pressure is somewhat more pronounced than the measured

influence, but the agreement is still satisfactory.

88

9 •

J) '7

e 8 .. 10 ~

Chapter 7 GENERAL DISCUSSION AND CONCLUSIONS

7.1. COMPARISON BETWEEN MODEL SLUDGE AND OTHER SLUDGES

As regards to the oxidation of glucose it was found that the·

results obtained were in fair agreement with the assumption

that the chemical reaction orders of glucose oxidation were

zero for oxygen and one for glucose.

In the case of oxidation of primary sludge it was found that

at oxygen partial pressures as applied in practice the chemi-~

cal reaction orders of the oxidation were equal to one for:·

both oxygen and organic matter.

Only at higher partial pressures of oxygen the chemical re­

action order for oxygen decreased to zero.

89

No special research was devoted to the reaction orders of ac­

tivated sludge. It was asswned that the orders were the same

as those of primary sludge. From an investigation of the in-·

fluence of the total pressure it was shown that this was a

good asswnption. The experimental results could be well de­

scribed with a mathematical model, starting from the above.

mentioned orders (see 6. S.) •

Under the same reaction conditions the model sludge of glucose

had a very high conversion rate, while the conversion rates of

activated and primary sludges were practically the same but

much lower.

A comparison of the physical absorption rate with the chemical

absorption rate of oxygen showed that the oxidation of glucose

was so fast that it took place inside the diffusion layer a­

round the air bubbles (see 3. 3. 1.) • As a consequence of this

diffusion limitation the partial chemical reaction orders of

zero in oxygen

into partial

oxidation of

and of one in organic material both converted

conversion orders of one half (see 3.3.3.). The

activated

slowly and conversion

and primary sludges proceeded more

took mainly place in the bulk of the

suspension.

The degree of diffusion limitation of oxygen or the extent to

which the conversion rate was reduced by oxygen transfer, de­

pended on temperature. At 180°c mass transfer limitation could

be neglected, while at 290°c the conversion rate was largely .

determined by the rate of mass transfer.

From the experimental results the reaction rate constants were

calculated. The rate constants for the oxidation of glucose ·

are presented in table 7.1.

90

Tab"le ( .. 1.

Reaation rate aonstant of g"luaose oxidation

temp k

(OC) (h-1)

190 2900

200 7500

These values were calculated from the continuous flow experi­

ments and are in good agreement ·with those calculated from the

semi-batch experiments. The oxidation of activated and primary sludge is rather more complicated since sludge does not consist of a single compound,

but of a diversity of products. It was found that the assump­

tion of simultaneous oxidation of two groups of components al­

lowed a fair description of the experimental results.

This was extensively investigated with the help of activated.

sludge (see 6.4.). About 65% of the activated sludge could be

considered to belong to the group of more reactive compounds·, 25% to the group of less reactive products, while about 10%

consisted of almost stable material. The reaction rate con­

stants for both groups are given in table 7.2, together with

the maximum conversion fmax' which is the fraction of oxidi­sable material.

In this table Isa and !sb are the reaction rate constants of the group of more reactive and of the group of less reactive com­

pounds, respectively. The effect of temperature on ~a and ~b could be described by Arrhenius'law. The activation energy for both groups was the same and had a value of 23'kcal/mol.

Experiments with primary sludge were carried out only at 230°c.

The reaction rate constant, which was based upon the total

amount of oxidisable matter, had a value of 17 m3/kg h (see

5.2.3.).

91

Tab Le ?. 2.

Reaation rate aonatante of aativated eiudge.

temp k ~b f max -a

{oC) {m3 /kg h) {m3/kg h) {%)

from from *) from from *)

experiments figure 6.5 experiments figure 6.6 . -·

180 2.9 2.3 0.35 0.30 90

200 9.0 9.0 0.97 0.97 90

220 24 23 3.6 3.4 90

230 39 4.0

240 43 58 7.9 6.2 91

255 180 130 12 12 92

270 250 250 20 22 93

290 500 500 22 50 93

*) Obtained by interpolation.

For a comparison of this value with the reaction rate constant

of activated sludge one has to take the product of ~a and the fraction of reactive compounds a • From table 7.2 it follows

3 0 that ~a = 39 m /kg h, so ~aao equals 25, as compared with a value of 17 of the reaction rate constant of primary sludge.

7 .,1. 4. Effect of dis111olution -------------~----~--

The effect of the dissolution of solid sludge particles on the conversion rate was studied using primary sludge.

From these investigations it was concluded that the dissolu­

tion did not influence the conversion rate since hydrolised

and solid sludge had nearly the same reaction rate constants.

So, for a description of the conversion rate the dissolution

step did not have to be taken into account {see 5.2.3.).

92

Experiments with a diversity of activated sludges showed that

the sludge produced at the sewage treatment plant of Eindhoven

which was used in most of the experiments, had almost the same

oxidation properties as the other sludges. Although this does.

not necessarily mean that ~a' ~band a0

always have the same.

values, it does mean that the conception of the two groups of

different reactivity can be applied also to other sludges than

that of Eindhoven.

7.2. ANOMALOUS PHENOMENA

Although the study of sludges of different origins showed that

activated sludges generally have the same oxidation proper-

ties, the activated (and also the

Eindhoven in the autwnn of 1969

it was before or afterwards. The

• primary) sludge produced at.

was much more reactive than

conversion rate in the· semi-

that the reaction took place mainly t'

batch reactor was so high

within the diffusion layer 0 ,.

around the gas bubbles. At 235 C

the initial conversion rate could be described with a reaction

rate constant of the order of 30,000 m3/kg h.

Since both activated and primary sludges had these high reac­

tion rate constants, it was assumed that they resulted from an

activator in the waste water of Eindhoven. It might be that

this activator had been discharged for some time by one of the

many galvanic industries of Eindhoven.

Analysis of the waste water of Eindhoven revealed that the

concentrations of copper, chromium, zinc, lead, nickel and

cadmium during this period were the. same as before and after

this period (126). So the activator could not be one of these

elements. From experiments with cobalt we found that it could

not be this element either.

This sludge had some more anomalous properties. One of these

was that when applying a large gas-liquid interfacial area,

resulting in a high c.onversion rate, fmax was 15-20% lower

93

than for a small interfacial area (low conversion rate) (121,

124) •

Another anomalous phenomenon was that when this sludge was ox­

idised in a continuously operated stirred tank reactor (=CSTR),

the high conversion rate was not observed while the reaction

rate constant was of the order of magnitude of "normal" sludge

(121, 123). Semi-batch experiments with mixtures of fresh

sludge and sludge oxidised in the CSTR did not show the high

conversion rate either. From this it was concluded that in the

CSTR compounds were generated which counteract the effect of

the activator (inhibition).

Semi-batch experiments were also carried out with mixtures of

fresh sludge and sludge oxidised in the semi-batch reactor

during several reaction times. Again, the high conversion rate

was observed. From this it might be concluded that the genera­

tion of the inhibitor was caused by the mixing, owing to which

all reaction products are present at the same time and place,

which was not the case in the semi-batch reactor. The genera­

tion of the inhibitor might proceed through a combination of

products which were not present at the same time in the semi­

batch reactor.

Although these phenomena are not yet understood,it may be con­

cluded that compounds exist which accelerate the conversion

rate considerably. When these compounds are present, the con­

tinuous oxidation should be carried out in a reactor with on­

ly a small spread in residence time (nearly plugf low) , since

mixing provides conditions in which products are generated

counteracting the effect of the activator.

7.3. CALCULATION OF THE SIZE OF COMMERCIAL REACTORS

In this section the size of a reactor for wet-air oxidation

is calculated, based upon the results of our research. It was

assumed that a reduction of the effective COD of 90% had to be

obtained. Since the maximum conversion is about 90%,the reduc­

tion of the total COD should be 81%.

94

Since the oxidation takes place in the bulk of the suspension

and is partly limited by the diffusion of oxygen, a gas bub­

ble reactor is most appropriate. If we assume uniform gas

bubbles, the gas phase may be considered to move in a plug

flow. The effect of the flow conditions of the liquid phase

on the desired holding time, has been determined for the fol­

lowing cases:

(i) The liquid phase is completely mixed.

(ii) The liquid rises in plug flow through the reactor, in

co-current with the gas bubbles.

(iii) The liquid moves downward in plug flow in counter-cur­

rent with the gas-bubbles.

The equations describing the three cases are derived from ox­

ygen balances. Since these procedures are analogous with those

set up in the chapters 5 and 6, just the resulting equations

and boundary conditions are given.

(i} The equations for this case are

d1T RT kQ,a ( 1T- p ) - dx =

H u e (7. 1} g

a. - a. = k Pf T a. p /H f e -a re e (7. 2)

13 - Be = ~b PfTr13ePe/H f ( 7. 3)

a.f + 13 - a. - 13 = uS Pf

u (l-1T ). f e e [cf ]RT Q, e ( 7. 4)

In these equations x is the length of the reactor and Tr

the residence time of the sludge phase in the reactor.

In chapter 3, M was introduced as the supply rate of ox­

idisable material over supply rate of oxygen, so

M =

95

96

Since [cf]= fmax o [cf], equation (7.4) can be trans­

formed into

f a + S - a - S = max (1 - TI )

f f e e zr- e (7.5)

With the boundary condition TI = l for x = 0,the solution

of equation (7.1) can be transformed into

l - p ln e =

TI -p e e (7.6)

in which X is the desired length of the reactor and Nt

the number of transfer units. Nt can be converted into

Substitution of the following characteristic values

M =

kta =

T = r

Pf/H =

Cc l = f

results in

N = 4 t

l

360 h -1

l h

0.2 kg/m3

20 kg/m3

Substitution of this in equation (7.6) shows that

So, the oxygen concentration in the liquid phase ap­

proaches the equilibrium concentration belonging to ox­

ygen pressure in the gas phase at the exit.

(ii) For the case · of co-current plug flow the simultaneous

differential equations are

dTI RT.k.R,a ( TI-p) = dx H.ug

- 2..e. k.R,a

(TI-p) <Jsa[co]a.p = - -- -dx UJ(,

The boundary conditions are

X = 0 TI = l

p = 0

a. = a.a Q = f _,.. µ max ... o.

(7.8}

+ Jsb[co]Sp)~-£ .!(,

(7.9)

(7.10)

(7.11)

owing to the exothermic oxidation an amount of heat is

generated which amounts to an increase of the tempera­

ture, which consequently results in an increase of the

amount of steam in the air flow.

Assuming that the air feed is saturated with steam, the

temperature rise 6t is given by the following heat bal­ance across the reactor

in which ¢.!(,and ¢g are the mass flow of sludge and air. \

respectively, cp is the heat capacity of the sludge, H:

is the humidity of the air in kg water per kg air at the indicated temperature (see figure 7.1), 6H is the heat

97

98

40 4

3 30 atm

water) I air

2

kg kg

1

0 220 240 260 280

temperature (0 c) Figure 7.1.

Humidity of air as funation of temperature and pressure.

of evaporation, 6COD is the decrease in the COD, p the

density of the sludge and 6r is the heat of reaction per

kg COD and ti is the inlet temperature.

From the data of Zimmermann (74) it follows that

6r ~ 3000 kcal/kg COD.

Assuming an average temperature

while [cf] is 20 kg/m3 and M = temperature rise 6t is 3°c at a

and 24°c at a total pressure of

inlet temperatures should then

spectively.

in the reactor of 240°c

1, it is found that the

total pressure of 40 atm

70 atm.The corresponding

be 239°c and 228°c,re-

For an exact solution of this case, the temperature pro­

file in the reactor has to be taken into account.

owing to the temperature rise the oxygen pressure will

decrease faster by evaporation of water than at con­

stant temperature, while on the other hand the reaction

rate constants will increase. At 70 atm and 22.8°c pf/H =· 3 0 3 0 ' 0.32 kg/m , at 240 C : 0.46 kg/m ·and at 252 C : 0.64. ·

As a first approximation it is assumed that both effects!

compensate each other.

(iii) For the counter-current plug flow the equations are.

nearly the same as those for the co-current plug flow •.

Only in equation (7.8) has the sign of the left hand

term to be changed from negative into positive.The boun-

•

· dary conditions are

x = 0

x = x 'JT = 1

p = 0

Ct = Ct 0

f3 = f -a max o

Also in this case the temperature of the sludge phase

will rise, but again it is assumed that the effect of

the decrease in the partial oxygen pressure caused by

evaporation is compensated for by the increased reaction

rate constants.

These three cases have been solved for a temperature of 24o0 c and pressures of 40 and 70 atm, with [cf] = 20 kg/m3 , 1 - c: = 0.95, kia = 0.1 sec-1 and M = 1 (stochiometric amount of oxy­

gen), M = 0.91 (10% surplus) and M = 0.77 (30% surplus).

The reaction rate constants were taken from the figures 6.5

and 6.6.

The calculated values of Tr' necessary to obtain a 90% reduc­

tion of the effective COD, are shown in table 7.3. It follows

99

Table 7.3.

Desired residence time in reaator.

time in minutes; temperature 240°c

40 atm 70 atm

M (-) 1.0 0.91 0.77 1.0 0.91 0.77

completely 1830 1086 592 496 234 137 mixed

co-current 450 320 216 91 69 47 plug flow

counter-current 144 130 117 31 28 25 plug flow

from the table that the counter-current plug flow provides the

smallest residence times and that by increasing surplus of ox­

ygen the differences between the three systems diminishes.

With respect to reactor volume the counter-current plug flow

reactor seems to be the most appropriate. However, in pract~ce

it will be difficult to create a good approximation of pure counter-current plug flow in such large diameter columns, b~­

cause of circulation (110, 116). It might be that the reaction can be carried out successfully in a plate column as is nor­

mally used for destillation purposes. The sludge flows over the plates downwards from plate to plate,

while the air moves from the bottom to the top of the column

through holes or bubble caps in the plates. Owing to the ne­

cessity of a gas space between the plates, the total reactor

volume will be about twice as large as followed from table 7.3.

So the reactor volume , and therefore also the reactor cost ,

will approach the cost of a co-qurrent reactor. Furthermore, the plates will cause the column to be very difficult of ac­

cess.

100

The use of packing will also provide the desired system, but

then the packing will take a large amount of reactor volume by

which the total reactor volume will approach that of the co­

current system. Furthermore fouling might occur, resulting in a choking up of the reactor.

From these considerations it can be inferred that even though

the counter-current plug flow reactor requires the shortest

residence time, in practice a co-current reactor will be most

appropriate.

Finally it can be concluded that the calculated required res­

idence time for such a reactor is of the order of magnitude of the residence time applied in practice.

101

REFERENCES

1. G.M. Fair, J.C. Geyer, D.A. Okun; Water and wastewater engineering, vol l & 2, Wiley, New York 1966 and 1968.

2. C.F. Gurnham; Principles of industrial waste treatment, Wiley, New York 1955.

3. R.J. Dyson; Proceedings of the second symposium on the treatment of waste waters, New Castle Upon Tyne, 1959, 265, Pergamon, Oxford 1960.

4. W.W. Eckenfelder Jr., D.J.0 1 Connor; Biological waste treatment, Pergamon, Oxford 1961.

5. W.W. Eckenfelder Jr., Industrial water pollution control, Mc Graw-Hill, New York 1966.

6. K. Imhoff; Taschenbuch der Stadt-Entwasserung, 22ed, Oldenbourg, Milnchen 1969.

7. J. Muskat; Chem.-Ing.-Techn. 1967, 39, 179.

8. E.P. Willimon, J.F. Andrews; J. Wat.Poll.Contr.Fed. 1969, !!_, 99.

9. G. Rohlich; Advances in water pollution research, vol. 2, 207, Pergamon, Oxford 1964.

10. B.W. Dickerson, F.J. Farrell; J.Wat.Poll.Contr.Fed.,1969, !!_, 56. .

11. L.K. Bos; Zuiveringsslib, herkomst en besternrning, Lit. Survey nr. 30, Centrum voor landbouwpublikaties en land­bouwdocumentatie, Wageningen 1966.

12. Rijksinstituut voor de zuivering van afvalwater, medede-ling nr. 6, 1968, Staatsuitgeverij, 's Gravenhage 1968.

13. A. Fischer; Chem.-Ing.-Techn. 1967, 39, 157.

14. Mc Carty; J.Wat.Poll.Contr.Fed. 1967, 39, 157.

15. H. Rilffer; Vom Wasser, 1967, .2_!, 345, Verlag Chemie, Weinheim 1968.

16. J. Griffiths; Proceedings of the second symposium on the treatment of waste waters, New Castle Upon Tyne, 1959,36~ Pergamon. Press, Oxford 1960.

17. J.D. Swanwick et al; Advances in water pollution research, vol. 2, 387, Pergamon, Oxford 1964.

18. H. Akeyl, M. Neven; Chem.-Ing.-Techn., 1967, 39, 172.

19. M.A. Kershaw; Proceedings of the second symposium on the treatment of waste waters, New Castle Upon Tyne, 1969,33~ Pergamon, Oxford 1960.

20. H.W. Marson; J. Inst.Sew.Puri£. 1965, (4), 320.

21. Brit.pat. 653.984.

102

22.

23.

24.

I.K.

R.B. G.H.

Porteous; Wat. and Waste Treatm. J., 1960, 7, 543.

Brooks; Wat.Poll.Contr., London, 1968, 67, 592.

~eletzke; Purdue Univ.Eng.Bull. ext.ser., 1965,118, 40.

25. G.H. Teletzke; Proc.Biochem., 1966, li 329.

26. T. Helfgott, P. Webber; Wat. Works ans Wastes Eng., 1965, _£, (9), 76.

27. F.J. Zimmermann; Ind.Wat. and Wastes, 1961, ~,102.

28. F.J. Angenend; Brennst.Warme-Kraft, 1966, ~' 252. 29. L. Reh; Chem.-Ing.-Techn., 1967, 12_, 165.

30. J.I. Frankel; Chem.Eng., 1966, 73, Aug. 29, 91.

31. E.S. Shannon; J.Wat.Poll.Contr.Fed., 1968, !Q_, 2058.

32. F.E. Dalton et al; J. Wat.Poll.Contr.Fed., 1968, 40, 789.

33. J.C. Wylie; Proceedings of the second symposium on the treatment of waste waters, New Castle Upon Tyne, 1959,349, Pergamon, Oxford 1960.

34. A.J. Steffen; Wat.and Wastes Eng., 1966, li (9), 63. 35. E.A. Pearson; Advances in water polution research, vol.3,

Pergamon, Oxford 1964.

36. H. Kayser; Umschau, 1969, .!.Q., 299.

37. J.M. Watson; Trans.Inst.Chem.Eng., 1969, 47, (The Chemi-cal Engineer), no 225, CE 4.

38. M.A. Kershaw: Proc.Biochem., 1969, li 46.

39. R.B. Dean, Environm.Sc. and Techn., 1968, _£, 1079. 40. R.B. Dean, Tappi, 1969, 52, 457. 41. German pat. 266096.

42. Swed. pat. 143765. 43. U.S. pat. 2.665.249.

44. U.S. pat. 2.773.026.

45. Swed. pat. 159.595.

46. Kirk-Othmer; Encycl.Chem.Techn., sec. ed. vol. 6, 992, Intersience, New York 1965.

47. U.S. pat. 2.824.085. 48. u.s. pat. 2.903.425.

49. U.S. pat. 2.932.613.

50. U.S. pat. 2.944.396. 51. U.S. pat. 3.060.118.

52. British pat. 706.686.

103

53.

54.

55.

56. 57.

58.

British

British

British I

British

British

British

pat. 785.267.

pat. 785.269.

pat. 809.604.

pat. 812.832.

pat. 1.054.239.

pat • 1.055.348.

. 59. British pat. 1.055.349.

60. German pat. 1.245.868.

61. Belgian pat. 677.038. 62. Dutch pat. 82309.

63. Dutch pat. 88965.

64. 65.

66.

67. 68.

69.

Dutch Dutch

Dutch Dutch Dutch

Dutch

pat. pat.

pat. pat. pat.

pat.

89125. 123415. appl. 6506648.

appl. 6507280. appl. 6508133. appl. 6701476.

70. F.J. Zimmermann; Chem.Eng., 1958, §..?_,August 25, 117.

71. S.M. Yunis; Ph.D.Thesis, Ill.Inst.Techn., Diss.Al:Jstr.1967, !:1 .. .t 3976.

72. F. Meinck, H. Stoof, H. Kohlschiitter; Industrie-Al:Jwasser, 528, Gustaf Fisher Verlag, Stuttgart 1968.

73. Notice in J.Wat.Poll.Contr.Fed., 1969, 41, research suppl. 2. ~

74. F.J. Zimmermann; Purdue Univ.Exptl.Sta.Eng. Bull. 1958, 96, 409 •

. 75. G.H. Teletzke et al; J.Wat.Poll.Contr.Fed., 1967, 39,994.

76. E. Hurwitz, G.H. Teletzke, W.B. Gitchel; Wat. and Sewage Works, 1965, Aug., 298.

77. Standard methods for the exam. of wat. and waste wat., Amer.Health Assoc. Inc., New York 1962.

78. H.A. Pray et al; Ind.Eng.Chem., 1952, !ii 1146.

79. E. Hurwitz, W.A. Dundas; J.Wat.Poll.Contr.Fed., 1960, 32, 918.

80. F. Abel et al; Sewage and Ind.Wastes, 1954, .£§_, 1450. 81. I.G. Peplyaev et al, CA 70 (1969), 31479.

82. A. Goldstein, S. Lokatz; paper presented at Wat.Poll. Contr. Fed. conference Seattle, 8 Oct. 1963.

104

83. W.R. Walters, G. Ettelt; Purdue Univ.Eng.Bull.Ext. Ser. 1965, 118, 551.

84. P.V. Danckwerts; Chem.Eng.Sci. 1953, £, 1.

85. E.T. v.d. Laan; Chem.Eng.Sci. 1957, 21 187.

86. G.H. Teletzke; Chem.Eng.Prog.1964, 60, (1), 33.

87. J. Muel~er et al; Biotechn. and Bio.Eng. 1968, 1.Q., 331.

88. J.B. Mc Kinley; Wat.Works and Wastes Eng. 1965, £, (9) ,97.

89. J.C. Harding, G.E. Griffin; J.Wat.Poll.Contr.Fed., 1965, 37, 1134.

90. G. Astarita, G. Marrucci; I.E.C. Fund. 1963, £, 4.

91. R.W. Nicholson et al; The American City, 1966, .§.!., 97.

92. F.J. Zimmermann, D.G. Diddams; Tappi 1960, ~' 710.

93. G.H. Teletzke; paper presented at 39th annual conference of Water Pollution Control Federation, Kansas City, 1966.

94. R.H. Bogan; J. Sani.Eng.Div., .1959, July, SA 4J 13.

95. E. Guccione; Chem.Eng., 1964, May 25, 118.

96. N. Seegert; Vattenhygien, 1965, £1 47.

97. E. Hurwitz; Advances in water pollution research, vol. 2, 411, Pergamon, Oxford 1964.

98. E. Schmidt; Ber.Abw. Techn.Verein., 1957, ~' 78.

99. H. Lohman, A. Tilly; Chem.-Ing.-Techn., 1965, 37, 913.

100. E.W. Schoeffel, N. Seegert; Wasser, Luft, Betrieb, 1966, 8, 541.

101. Stauffer; Aufbereitungstechnik, 1967, ~' 286.

102. R. Randolf; Wasserwirtsch.-Wassertechn., 1963, 8, 10,342.

103. s. Lerczynski; Przem.Chem., 1967, 46, 2.

104. L. Gromski; Przem.Chem. 1967, !§.i_ 191.

105. R. Vrijburg; Water, 1966, 50, 313.

106. B. Brunes et al; Svensk Papperstidning, 1955, 58, 332.

107. K.N. Cederquist; Svensk Papperstidning, 1958, .§.!, 38.

108. K.N. Cederquist; Svensk Papperstidning, 1955, 58, 154.

109. K.N. Cederquist, P. Bering; Acta.Polytechn.Chem.Met.Ser. 1952, .~.' 5.

110. G.D. Towell et al ; A.I.Ch.E., Chem. Eng. Symp. Series

no • 1 O , 19 6 5 . 111. R.S. Burd; Water pollution control research series,

publ. WP-20-4, A study of sludge handling and disposal, 275.

112. U.S. Patent 3.449.247.

105

113. P. Battino; Chem.Rev. 1966,.§.§_, 395.

114. Dutch pat. appl. 6917974.

115. K.M. Guthrie; Chem. Engin. 1969,~ March 24, 114.

116. J-B. Wijffels; Ph.D. Thesis, University of Technology, Eindhoven 1970.

117. R.H. Perry; Chemical Engineers Handbook, fourth edition, McGraw-Hill, New York 1963.

118. R. Higbie; Trans. A.I.Ch.E. 1935, l!_, 365.

119. R.B. Brooks; Wat.Poll.Contr. London, 1970, .§..2_, 92.

120. Takeichiro Takamatsu,et al, Wat. Research, 1970, ii 33.

121. J.J.A. Ploos van Amstel; Internal publication, University of Technology, Eindhoven 1970.

122. H.G.J. de Wilt; Ph.D. Thesis, University of Technology, Eindhoven 1969.

123. C.G.M. de Boer; Internal publication, University of Technology, Eindhoven 1970.

124. H.P.E. v.d. Venne, F.C.M. v.d. Berg; Internal publication University of Technology, Eindhoven 1970.

125. J.J.A. Ploos van Amstel, K. Rietema; Chem. - Ing. - Techn. 1970, !£, 981.

126. H.J. Eggink, Private communication.

127. J.V. Hunter, H. Heukelekian; J. Wat.Poll.Contr.Fed. 1965,

]]_, 1142. 128. u. Grigull; Die grundgesetze der WarmeUbertragung, dritte

auflage, Springer 1963.

129. T. Takamatsu et al; Paper II-32 presented at the 5th

international Water Pollution Research Conference 1970.

106

NOMENCLATURE

[a]

a

af [b]

[cb]

[cf]

[ci]

[ce]

[co]

[cd]

[cs]

[ct]

[e]

[emaxl

(eo]

Dr

D 0

db

df E

f

f max

G

H

COD of more reactive matter

specific gas-liquid interface

specific outer surf ace of flocks

COD of less reactive matter effective COD in bulk

effective COD in feed

effective COD at interface

effective COD in exit

effective COD at time zero

effective COD of dissolved matter

effective COD of suspended sludge

effective COD of suspension

COD of sludge

solubility of sludge, in COD

initial COD of suspension

diffusivity of dissolved organic material

diffusivity of oxygen

average bubble diameter

average flock diameter

eddy diffusivity

conversion

maximum conversion

gas flow rate at reaction conditions

Henry coefficient

kg/m3

-1 m -1 m

kg/m3

kg/m3

kg/m3

kg/m3

kg/m3

kg/m3

kg/m3

kg/m3

kg/m3

kg/m3

kg/m3

kg/m3

m2/sec 2; . m sec

m

m

m2/sec

%

%

m3/sec atm m3

kg

k reaction rate constant sec-1 ,h-l

k oxidation rate constant of reactive matter m3/kg h · -a ~b oxidation rate constant of less reactive matter m3/kg h

oxidation rate constant of sludge suspension

oxidation rate constant of dissolved matter

hydrolysis rate constant

liquid side mass transfer coefficient at g-1 interface m/sec

107

M

N r

[ 0]

[ob]

[oi]

[of]

p

Pf

Pe

Pe

x

x

108

liquid side mass transfer coefficient outside flocks

constants in sludge oxidation model

Liquid flow rate at reaction conditions

u~ RT [cf] _ supply rate of effective COD ugpf - supply rate of oxygen

m/sec -1

h -1

h

m3/sec

V 2 D k RT o- aa X(overall number of conversion H u~ug N stages)

oxygen concentration

oxygen concentration in bulk

oxygen concentration at g-1 Pf H oxygen partial pressure

oxygen partial pressure in

oxygen partial pressure in

UR, X

E Peclet number

average oxygen flux through

oxidation rate

oxidation rate of dissolved

interface

feed

exit

g-1 interface

matter

oxidation rate of suspended particles

oxidation rate of suspension

gas constant

time

absolute temperature

superficial gas velocity

superficial liquid velocity

reactor volume

distance in reactor

length of reactor

kg/m3

kg/m3

kg/m3

3 kg/m

atm

atm

atm

kg/m2h

kg/m3h

kg/m3h

kg/m3h

kg/m3h

atm m3

OK kg 02

sec,min,h

OK m/sec

m/sec m3

m

m

z

a 0

s y

y*

-y

e:

p

a

T

e

<Pox

distance from interface

[a]/[e0

]

fraction of reactive matter in sludge

[b]/[c0

]

[c]/[c0

]

[c]/[cf] [c]/[e

0]

gas fraction in reactor

volume fraction of flocks in suspension

volume fraction solids in flock

p/pf

[ob]/[of] x/X

m

residence time of liquid element at g-1 interface sec

average residence time of sludge in reactor sec,min

conversion rate

109

SAMENVATTING

Natte oxidatie is een methode van slibverwerking waarbij het

·slib in de aanwezigheid van lucht in de vloeibare waterfase

geoxideerd wordt bij temperaturen van 200-300°c en drukken van

40-120 atm.Uit een literatuurstudie bleek dat dit proces langs

empirische weg was ontwikkeld en dat inzicht in de achtergrond

van de natte oxidatie nauwelijks aanwezig was •

Vanwege de aantrekkelijkheid van het proces werd een proces­

kundig gericht onderzoek uitgevoerd, hetgeen in dit proef­

schrift beschreven is.

Allereerst werd de oxidatie van een glucose oplossing als mo­

del slib onderzocht met semi-batch en continue proeven bij om­

streeks 200°c en 50 atm. Uit deze experimenten bleek dat de

oxidatie van het model slib snel was vergeleken met de diffu­

sie van zuurstof, waardoor de reaktie zich in de diffusie­

grenslaag rond de gasbellen afspeelt. De chemische reaktie

snelheid kan beschreven worden als eerste orde in organisch

materiaal en nulde orde in zuurstof, terwijl de reaktie snel­

heidskonstante ongeveer 2 sec-l is bij 200°c. Een model wordt

gepresenteerd voor de conversie in de continue reaktor, dat

zowel gekombineerde reaktie en diffusie inhoudt, als menging

en konvektie.

Dit model geeft een akseptabele beschrijving van zowel de in­

vloed van de proces parameters als van de koncentratieprof ie­

len in de reaktor.

Bij de verhoogde temperaturen (ca 240°c) gaat het slib ten ge­

volge van hydrolyse gedeeltelijk in oplossing, waardoor zowel

oxidatie van gehydroliseerd slib als van slib deeltjes op­

treedt.

Om de bijdrage van de oxidatie van gehydroliseerd slib tot de

totale oxidatie snelheid te kennen werd de hydrolyse van slib

met batch experimenten onderzocht.

110

Het effect van de hydrolyse op de . overall oxidatie snelheid

werd bestudeerd aan de hand van primair slib. Oxidatie experi­

menten werden uitgevoerd bij 230°c en 100 atm met slib als zo­

danig met oplossingen van gehydroliseerd slib en met een sus­

pensie van het vaste residu dat na de hydrolyse overblijft.

Het blijkt dat de hydrolyse de overall oxidatie snelheid niet

beinvloedt omdat gehydroliseerd slib en slibdeeltjes vrijwel

dezelfde reaktiviteit hebben.

De oxidatie van aktief en primair slib verliep langzamer dan

de oxidatie van het model slib, en de omzetting vond voorna­

melijk in de bulk van het slib plaats. De mate van diffusie­

limitering door zuurstof, ofwel de mate waarin de omzettings­

snelheid door zuurstoftransport wordt afgeremd, hangt af van ·de temperatuur. Bij 180°c kan de weerstand voor massa trans-port verwaarloosd worden, terwijl bij 290°c de omzettings­

snelheid grotendeels door massa transport bepaald wordt.

Een model voor de oxidatie van slib wordt gepresenteerd. Het

uitgangspunt van dit model is dat in het slib twee groepen van

komponenten onderscheiden kunnen worden die verschillen in re­aktivi tei t en die simultaan geoxideerd worden, terwijl nog een

derde, volledig inactieve groep, eveneens aanwezig is.

Experimenteel is vastgesteld dat aktief slib voor ca 65% uit

reaktief materiaal bestaat, voor 25% uit minder reaktief ma­

teriaal terwijl 10% volledig inaktief is. De chemische reak­

tiesnelheid wordt beschreven als eerste orde in organisch ma­

teriaal en als eerste orde in zuurstof (bij relatief hoge zuur­stof drukken wordt de reaktie nulde orde in zuurstof).

Massa transport is ook in het model begrepen. Dit model geeft

een zeer akseptabele beschrijving van de experimentele resul­

taten. Het effect van de temperatuur op de snelheidskonstan-

ten van de twee groepen van oxideerbare komponenten

slib kan met de wet van Arrhenius beschreven worden, de aktiveringsenergie voor beide groepen ongeveer 23

bedraagt.

in aktief

terwijl

kcal/mol

111

Bij 2ss0 c ziJn de reaktie snelheids konstanten respectieve­

lijk 250 en 20 m3/kg h.

Gebaseerd op de resultaten van dit onderzoek werden de afme­

tingen van een kommerciele sektor berekend. Afhankelijk van de

overmaat zuurstof, is de benodigde verblijftijd voor een prop­

stroom reaktor, bedreven in gelijkstroom, twee tot drie maal

groter dan in een in tegenstroom bedreven propstroom reaktor.

De benodigde verblijftijd in een reaktor met een ideaal ge­

mengde vloeistof f ase is 6 tot 12 maal groter dan in de tegen­

stroom reaktor. In de praktijk wordt een gelijkstroom reaktor

toegepast waarin de mengtoestand tussen propstroom en ideaal

gemengd ligt, waardoor een verblijftijd van ca 5 maal de ver­

blijftijd in een in tegenstroom bedreven propstroom reaktor

nodig zal zijn. Op praktische gronden zal echter de gelijk­

stroom reaktor waarschijnlijk toch aantrekkelijker blijven.

112

LEVENSBESCHRIJVING

De schrijver van dit proefschrift werd in oktober 1943

te Eindhoven geboren.

Na het behalen van het HBS-B diploma in 1961 ving hij aan met

zijn studie voor scheikundig ingenieur aan de Technische Hoge­

school te Eindhoven. Vanaf 1962 tot aan het einde van zijn

studie was hij student-assistent op het laboratorium voor al­

gemene chemie en op het laboratorium voor fysische technologie.

Na een afstudeerperiode onder leiding van Prof .Dr. K. Rietema,

hoogleraar in de fysische technologie, werd de studie in mei

1967 afgesloten. In mei 1967 vond tevens zijn benoeming plaats

als wetenschappelijk medewerker aan het laboratorium voor

fysische technologie, waar het in dit proefschrift beschreven

onderzoek verricht werd.

Eind 1970 trad hij als milieuhygienist in dienst van het ar­

chi tecten en ingenieursbureau van de N.V. Philips'Gloeilampen­

fabrieken.

113

STELLINGEN

1. De mono-nitrering van tolueen in het twee-fasen systeem

tolueen-nitreerzuur verloopt bij lage zwavelzuur concen­

traties hoofdzakelijk in de organische fase, terwijl bij

hogere zwavelzuur concentraties de reaktie zich hoofdzake­

lijk in de zuurfase afspeelt.

K.J. Jacobs; Afstudeer rapport, T.H. Eindhoven 1966.

P. van Galen; Afstudeer rapport, T.H. Eindhoven 1969.

2. Bij de mono-nitrering van tolueen in de organische fase

(lage zwavelzuur concentraties) wordt het aktieve nitre­

rings agens alleen in de zuurfase gevormd.

Deze nitrerings agens ontleedt onder invloed van water.

3. Voor het stoftransport tussen gas- en vloeistoffase in een

trickle kolom kan voor laminaire stromings condities de

volgende relatie afgeleid worden voor de weerstand in de

vloeistoffase.

Sh = 0.63 Re 1 1 3Go 1 16 sc~

4. Voor het digitaal of analoog oplossen van concentratie­

verdelingen in procesapparatuur waarin de mengtoestand ge­

karakteriseerd wordt door propstroom met axiale menging,

biedt het ten aanzien van de stabiliteit van de oplossing

grote voordelen om de ingangsrandvoorwaarde te vervangen

door een "overall" massabalans.

5. Yunis, die de oxidatie van glucose bij ca iso0 c bestudeerd

heeft, veronderstelt conversiesnelheden gemeten te hebben.

Hij heeft echter slechts de maximale conversie gemeten.

Zijn konklusies zijn dan ook onjuist.

M.S. iunis; Thesis, Ill. Inst. of Techn. Chicago 1967.

6. Het is zeer twijfelachtig of het model dat Takamatsu geeft

voor de hydrolyse van aktief slib bij verhoogde temperatuur

met de werkelijkheid overeen stemt. Het hierop gebaseerde model voor de natte oxidatie is niet in overeenstemming met het model dat in dit proefschrift gepresenteerd wordt.

T. Takamatsu et al; Wat. Res. 1970, ii 33. T. Takamatsu et al; Paper II - 32 presented at the 5th

international Water Pollution Research Conference, 1970.

7. De verontreinigingen in het afvalwater van de aardappel­

meelindustrie zijn goed te oxideren in de vloeibare water­fase bij verhoogde temperatuur en druk. De reaktiesnel­

heids konstanten zijn ongeveer een faktor 10 groter dan

die voor aktief slib.

J. Oomen enc. Trentelman; Praktikum verslag,

T.H. Eindhoven, 1970.

8. Schieten op het wipje zonder het jobje getuigt van de ware

schuttersgeest.

J.A. Jolles; De Schuttersgilden in Noord-Brabant

Handboek Hinderwet XXII - 4; oktober 1970.

9. Het is onduidelijk waarom bij het schieten op de vogel al­

leen het gebruik van rookzwak buskruit is toegestaan.

Handboek Hinderwet XXII - 7; oktober 1970.

10. Tot onze kultuur behoort karnaval, het feest waarbij af­

stand genomen wordt van de komplikaties van onze kultuur.

Nuenen, 2 april 1971. J.J.A. Ploos van Amstel.