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Chemical Engineering Communications

ISSN: 0098-6445 (Print) 1563-5201 (Online) Journal homepage: http://www.tandfonline.com/loi/gcec20

Effect of Various Curved-Blade ImpellerGeometries on Drop Size in a Liquid–Liquid StirredVessel

Reza Afshar Ghotli, A. R. Abdul Aziz & Shaliza Ibrahim

To cite this article: Reza Afshar Ghotli, A. R. Abdul Aziz & Shaliza Ibrahim (2017): Effect ofVarious Curved-Blade Impeller Geometries on Drop Size in a Liquid–Liquid Stirred Vessel,Chemical Engineering Communications, DOI: 10.1080/00986445.2017.1325738

To link to this article: http://dx.doi.org/10.1080/00986445.2017.1325738

Accepted author version posted online: 05May 2017.Published online: 05 May 2017.

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Effect of Various Curved-Blade Impeller Geometries on Drop Size in a Liquid–Liquid Stirred Vessel

REZA AFSHAR GHOTLI, A. R. ABDUL AZIZ, and SHALIZA IBRAHIM

Department of Chemical Engineering, Faculty of Engineering, University of Malaya, Kuala Lumpur, Malaysia

An experimental study was performed to discuss the effects of curvature angles and central disk sizes of 6-curved-blade impellers on the mean drop size in an agitated vessel. A system with 1%�oil in water in the presence of a surfactant solution was used. The effects of impeller speed on drop size were also investigated. The laser diffraction technique and RSM method were employed to measure and analyze data, respectively. Decreased curvature angles from 180° to 140° reduced the drop size up to 9%, 10%, and 10%�at agitation speeds of 5, 6, and 7 rps, respectively. Moreover, a decrease in central disk size from 3/4D to 1/4D reduced the drop size up to 16%, 18%, and 22%�at an agitation speed of 5, 6, and 7 rps, correspondingly. Two mathematical models were suggested and the most significant parameters of each experimental design were identified through the Analysis of Variance.

Keywords: 6-Curved blade impeller; Central disk size; Curvature angle; Liquid–liquid mixing; RSM; Sauter means diameter

Introduction Liquid–liquid mixing is a key stage for various applications in chemical, pharmaceutical, petroleum, and food manufacturing processes. Examples of industrial processes involving liquid– liquid mixing include polymerization, emulsification, and solvent extraction. In all these processes, drop size distribution is among the most important parameters for evaluating disper-sion stability and efficiency of operation. It plays an important role in affecting the mass transfer rate between the phases in liquid–liquid systems (Giapos et al., 2005; Maaß et al., 2007). Smaller drop sizes are more advantageous in mass transfer processes as they produce larger interfacial and mass transfer areas around the impeller zone compared to drops with larger size (Abdul Aziz et al., 2007; Patil and Kumar, 2010).

Performance characterization of a liquid–liquid agitation system is generally performed under dynamically equilibrium conditions between drop break-up and coalescence (Zaldívar et al., 1996; Sechremeli et al., 2006). These conditions can be created with the use of a surfactant in the experiments (El-Hamouz, 2009). Drop size distribution is affected by inversion between drop break-up and coalescence. Figure 1 shows the schematic illustration of drop break-up and coalesc-ence (Carlucci, 2010). The exact mechanism of coalescence and break-up in a liquid–liquid agitation system is extremely

complex and has not been well understood (Zaldívar et al., 1996). Drops break-up generally occurs in zones of high shear stress near the impeller blades depending on the mixing inten-sity Kraume et al., 2004). Moreover, turbulent velocity, viscous friction, and pressure variations also tend to break drops while collision between droplets may consequently cause coalescence (Kumar, 1983; El-Hamouz, 2009). Therefore, it is evident that dispersion with uniform drop size in the mixing process is not achievable (El-Hamouz, 2009).

Drop size distribution in liquid–liquid agitated systems has been the subject of previous studies. Different models and theories have been established to predict drop size distribution and the mean drop size based on different parameters. Kolmogorov’s local isotropy theory (Kolmogorov, 1949) is commonly used to evaluate drops break-up and drop size distributions. This theory proposes that break-up occurs when the inertial stress is greater than the interfa-cial tension stress (Zaldívar et al., 1996; Nienow, 2004). Hinze (1955) was the first to express a model based on Kolmogorov’s theory and show that drops in the inertial region of turbulence had the maximum size (dmax) (El-Hamouz et al., 2009). Both Hinze’s and Kolmogorov’s theories assume equilibrium conditions and indicate that there is a correlation between the maximum drop size diameter and dimensionless Weber number (Calabrese et al., 1986a; Pacek et al., 1999).

dmax

Da e� 0:4

Ta We� 0:6 ð1Þ

where D is the impeller diameter, eT is the energy dissipation rate, and We is Weber number which is determined by:

We ¼qcN2D3

rð2Þ

none defined

Address correspondence to A. R. Abdul Aziz, Department of Chemical Engineering, Faculty of Engineering, University of Malaya, 50603 Kuala Lumpur, Malaysia. E-mail: azizraman@ um.edu.my

Color versions of one or more of the figures in the article can be found online at www.tandfonline.com/gcec.

Chemical Engineering Communications, 0:1–13, 2017 Copyright © Taylor & Francis Group, LLC ISSN: 0098-6445 print/1563-5201 online DOI: 10.1080/00986445.2017.1325738

The average energy dissipation rate for each experiment was estimated based on the volume of the mixing tank (Giapos et al., 2005; O’Rourke and MacLoughlin, 2005; Sechremeli et al., 2006):

e ¼P

qVTð3Þ

P ¼ P0qN3D5 ð4Þ

where P is the power consumption, q is the density of the liquid, and VT is the liquid volume in the vessel.

The average drop size, which is called Sauter mean diameter, d32, is usually applied to evaluate drop size.

d32 ¼

Pnid3

iPnid2

ið5Þ

where, ni and di show the number and nominal diameter of drops, respectively. The average drop size is mostly used to cal-culate the interfacial area and many researchers have attempted to relate d32 to dmax (Lovick et al., 2005). Most of the published experimental works have reported that the maximum drop size is proportional to d32 due to the linear correlation between them (Sprow, 1967; Calabrese et al., 1986a; Zhou and Kresta, 1998a; Sis et al., 2005; El-Hamouz et al., 2009);

d32

D¼ C1We� 0:6 ð6Þ

where D is the impeller diameter and C1 is the dimensionless constant obtained experimentally depending on the tank geometry and impeller type (Podgórska, 2009). This equation has been verified for low dispersed phase volume fraction. Therefore, the rate of coalescence can be neglected (Pacek et al., 1999; Podgórska, 2009).

The mean drop size and drop size distribution in agitation vessels strongly depend on the geometry and impeller design (Paul et al., 2004; El-Hamouz, 2009; El-Hamouz et al., 2009; Podgórska, 2009). Impellers of high energy input have high energy dissipation rate per mass and therefore there is increased break-up rate or reduced coalescence rate (Giapos et al., 2005). Impellers that produce radial flows such as Rushton turbine, which is an impeller with high shear rate, is a common choice for a wide range of mixing applications, especially for gas dispersion and drop break-up applications (Oldshue, 1986; Pacek et al., 1999; Afshar Ghotli et al., 2013). Hydrofoil impel-lers such as Prochem Maxflow T, Lightnin A315 or Chemineer HE3 can provide rigorous axial flow with low power consump-tion (Podgórska, 2009). Several works on the effect of physical properties and geometry of agitation systems on drop size distri-bution have been published. However, the literature shows that

most of the experimental works have been conducted with Rushton turbines while the other types of impellers have rarely been used Fernandes and Sharma (1967) observed the same interfacial area for paddles with straight, inclined and curved blades. Thus, it can be concluded that the average value of drop size for all the studied impellers was the same. Furthermore, their results showed that greater area was provided by a six-bladed disk turbine compared to an open-style impeller (Sechremeli et al., 2006). Brown and Pitt (1974) proved that width of impeller blade did not have any significant effect on drop size distribution at the same impeller diameter and agitation speed (Giapos et al., 2005; Sechremeli et al., 2006). Zhou and Kresta (1998b) conducted a drop size study in a very dilute liquid–liquid system with a Rushton turbine and three axial flow impellers; they did not make any comparison between the impellers (Zhou and Kresta, 1998b). Pacek et al. (1999) reported the drop size distribution for Rushton turbine, disk turbine and four Chemineer impellers with both viscous and non-viscous dispersed phases by volume fraction of 1 and 5%. The results showed that low-power impellers produced the same drop size at the same mean specific energy dissipation rate. The results also proved that the drop size was much smaller for these impellers than that of a Rushton turbine and disk turbine (Pacek et al., 1999). Musgrove et al. (2000) examined the drop size distribution of immiscible liquid dispersions for Rushton turbines, pitched-blade turbines and hydrofoils including Lightnin A310, Chemineer and HE3. The results implied that hydrofoil impellers produced smaller average drop size (d32) than turbine impellers at the same power per unit volume and impeller diameter (Musgrove et al., 2000). Some of the previous works on drop size distribution are summarized in Table I.

In the current work, data on drop size measurement for dispersion of oil in water using various designs of 6-curved- blade impellers were obtained. The effect of curvature angles, central disk sizes and agitation speed on the mean drop size (d32) was investigated. An experimental design technique was also applied to develop statistical models to show the interac-tions between the parameters. Furthermore, the importance of the affecting parameters was determined via the Response Surface Methodology (RSM) (Alam et al., 2007; Bezerra et al., 2008; Arami-Niya et al., 2012; Afshar Ghotli et al., 2013). There is no information on the modeling of the interac-tion between angles, central disk size and speed with Sauter Mean Diameter (d32) in the existing literature. Therefore, in the present work, Central Composite Design (CCD) based on the Response Surface Methodology (RSM) was applied to study the effects of curvature angles of blades and central disk sizes on the mean drop size (d32) at different agitation speeds in a stir-red vessel and develop a model based on the experimental data.

Materials and Method

Materials Palm oil utilized in the present work was purchased from Sik Cheong Edible Oil Sdn. Bhd., Malaysia. The sodium dodecyl sulfate (SDS) in powder form was used as a surfactant stabilizer. The analytical-grade surfactant with a purity of >99.5%� was

Fig. 1. Schematic representation of drop break-up and coalescence in the liquid-liquid system.

2 R. A. Ghotli et al.

Tabl

e I.

Som

e of

the

typi

cal w

ork

cond

ition

s an

d m

easu

rem

ent m

etho

ds

Tank

Im

pelle

r ty

pe

Syst

em

Dis

pers

ed

phas

e %

�Su

rfac

tant

M

etho

d R

efer

ence

St

anda

rd s

tirre

d

vess

el

−6-b

lade

turb

ine

Styr

ene/

wat

er

1 po

lyvi

ny1

al

coho

l (PV

A)

LDM

C

hatz

i et a

l. (1

991)

Stan

dard

stir

red

ve

ssel

-R

usht

on tu

rbin

e W

ater

/org

anic

pha

se

0.8

Sodi

um d

odec

yl

sulfa

te (

SDS)

PM

Sa

thya

gal

et a

l. (1

996)

St

anda

rd s

tirre

d

vess

el

-Rus

hton

turb

ine

Dis

tille

d w

ater

/xyl

ene

10 to

80

0.3%

�sod

ium

dod

ecyl

su

lfate

(SD

S)

LDM

B

oye

et a

l. (1

996)

Stan

dard

stir

red

ve

ssel

-R

usht

on tu

rbin

e -H

E3

-Pitc

hed

blad

e tu

rbin

e -A

310

DIU

F w

ater

/sili

con

oil

0.03

N

ot u

sed

LDM

Zh

ou a

nd

Kre

sta

(199

8b)

Stan

dard

stir

red

ve

ssel

-R

usht

on tu

rbin

e -6

-bla

de im

pelle

r (6

DT)

-H

E3S

-HE3

L -C

S2

De-

ioni

zed

wat

er/c

hlor

oben

zene

an

d Su

n flo

wer

oil

1 an

d 5

Not

use

d V

MS

Pace

k et

al.

(1

999)

Stan

dard

stir

red

ve

ssel

-L

ight

nin,

A31

0

-Che

min

eer,

HE3

-R

usht

on tu

rbin

e -p

itche

d bl

ade

turb

ine

-Sili

cone

oil

-Chl

orob

enze

ne

-Xyl

ene

-Cyc

lohe

xane

-T

ri-bu

tyl p

hosp

hate

0.13

N

ot u

sed

PM (

Mac

ro

phot

ogra

phy)

M

usgr

ove

et

al.

(200

0)

Stan

dard

stir

red

ve

ssel

-R

usht

on tu

rbin

e -P

addl

e im

pelle

r To

luen

e-w

ater

1

and

10

Not

use

d V

MS

Rib

eiro

et a

l. (2

004)

Stan

dard

stir

red

ve

ssel

-R

usht

on tu

rbin

e D

e-io

nize

d w

ater

/sili

cone

oi

ls (

Dow

Cor

ning

200

) 0.

3 So

dium

dod

ecyl

su

lfate

(SD

S)

PVM

O

’Rou

rke

and

M

acLo

ughl

in (

2005

) St

anda

rd s

tirre

d

vess

el

−6-b

lade

turb

ine-

type

stir

rer

Dod

ecan

e/di

still

ed w

ater

0.

1 N

ot u

sed

LDM

Si

s et

al.

(200

5)

Stan

dard

stir

red

ve

ssel

-R

usht

on tu

rbin

e K

eros

en/w

ater

5

and

10

Not

use

d O

RM

Lo

vick

et a

l. (2

005)

Stan

dard

stir

red

ve

ssel

Fl

at v

ertic

al d

isk

styl

e Im

pelle

rs w

ith 2

,4,6

,8 b

lade

s K

eros

en/w

ater

1,

2.5

and

7 N

ot u

sed

PM

(pho

tom

icro

grap

hy)

Gia

pos

et a

l. (2

005)

Stan

dard

stir

red

ve

ssel

-R

usht

on tu

rbin

e H

eavy

oil

in T

olue

ne

25

Not

use

d LD

M

Ang

le a

nd

Ham

za (

2006

) St

anda

rd s

tirre

d

vess

el

−6-b

lade

dis

k st

yle

-6-b

lade

ope

n st

yle

Dis

tille

d w

ater

and

Ker

osen

e 1,

2.5,

5 an

d 10

N

ot u

sed

PM

(Pho

tom

icro

grap

hy)

Sech

rem

eli

et a

l. (2

006)

Es

co m

ixin

g ta

nk

-Hig

h sh

ear

Saw

toot

h Si

licon

oil

(Dow

co

rnin

g 20

0) /w

ater

1

Sodi

um la

uret

h

sulfa

te (

SLES

) LD

M

El-H

amou

z (2

007)

Esco

mix

ing

tank

-P

itche

d bl

ade

turb

ine

-Saw

toot

h Si

licon

oil

(Dow

co

rnin

g 20

0) /w

ater

1

Sodi

um la

uret

h

sulfa

te (

SLES

) LD

M

El-H

amou

z

et a

l. (2

009)

St

anda

rd s

tirre

d

vess

el

-Rus

hton

turb

ine

Tolu

ene/

wat

er

20

Not

Use

d −2

D O

RM

-F

BR

-F

BR

M

-ET

Maa

ß et

al.

(2

010)

(Con

tinue

d)

3

Tabl

e I.

Con

tinue

d

Tank

Im

pelle

r ty

pe

Syst

em

Dis

pers

ed

phas

e %

�Su

rfac

tant

M

etho

d R

efer

ence

St

anda

rd s

tirre

d

vess

el

−6-B

lade

Rus

hton

turb

ine

Suga

r sy

rup

/sun

flow

er o

il 3

Sodi

um d

odec

yl

sulfa

te (

SDS)

-L

DM

Pa

til a

nd

Kum

ar, (

2010

) M

ixin

g ce

ll −6

-Bla

de tu

rbin

e W

ater

-in-c

rude

oil

emul

sion

s 15

wat

er

Nat

ural

sur

fact

ant

prop

ertie

s of

the

cr

ude

oils

-FB

RM

-P

VM

B

oxal

l et

al.

(201

1)

Stan

dard

stir

red

ve

ssel

-R

etre

at c

urve

Im

pelle

rs (

RC

I)

Dis

tille

d w

ater

/ani

sole

, cyc

lohe

xane

, N

-but

yl c

hlor

ide

and

Tolu

ene

5, 2

0 33

, and

45

Poly

vin

yl

alco

hol (

PVA

) ET

M

aaß

et

al.

(201

2)

Bat

ch r

eact

or

−6-b

lade

pad

dle

-LR

4 Si

lver

son

ro

tor–

stat

or m

ixer

Aqu

eous

sol

utio

n of

N

aOH

/sol

utio

n of

ben

zoic

aci

d

and

ethy

l chl

oroa

ceta

te

in to

luen

e

1 N

ot u

sed

LDM

Ja

sińs

ka

et a

l. (2

012)

Stan

dard

stir

red

ve

ssel

6-

blad

e R

usht

on T

urbi

ne

Wat

er/s

ilico

ne o

ils

(20,

350

, and

500

mPa

s)

1 So

dium

dod

ecyl

su

lfate

(SD

S)

LDM

Za

inal

Abi

din

et

al.

(201

4)

LDM

: las

er d

iffra

ctio

n m

etho

d (p

artic

le s

ize

anal

yzer

). PM

: pho

togr

aph

met

hod.

O

RM

: opt

ical

ref

lect

ance

mea

sure

men

t. PV

M: p

artic

le v

ideo

mic

rosc

ope

prob

e.

FBR

M: f

ocus

bea

m r

efle

ctan

ce m

easu

rem

ent.

FBR

: for

war

d–ba

ckw

ard

ratio

sen

sor.

VM

S: v

ideo

mic

rosc

ope

syst

em.

ET: e

ndos

cope

tech

niqu

e.

4

supplied by Merck Chemicals Co., Germany. Palm oil with a vis-cosity of 0.08198 kg/ms and density of 890.00 kg/m3 was used as the dispersed phase. Table II shows the physical properties of the dispersed and continuous phases in the experiments.

Experimental Setup A transparent flat-bottom scratch-proof Perspex tank with a diameter of 0.4 m (T) was used as the reaction vessel. The vessel was equipped with four equally spaced wall-mounted baffles (B) with a width of B ¼ T/10. The ratio of impeller clearance (C) to tank diameter (T) followed the standard geometries and was equivalent to 0.133 m. Based on the objectives of this project, five types of 6-curved-blade impellers with various central disk sizes and curvature angles were used. The structure and descrip-tion of each impeller are given in Table III. The power number values in this work were obtained from the work published by Afshar Ghotli et al. (2013). All the power consumption measurement was performed using suspended load measuring system. The schematic illustrations of the experimental setup and impellers are presented in Figures 2 and 3.

Experimental Procedure In order to investigate the effects of different designs of 6-curved-blade impellers on Sauter Mean Diameter (d32) and drop size distribution, a series of experiments were carried out at atmospheric pressure and room temperature in a range of 297 K to 298 K. Tap water was used directly as the continuous phase in the system and the liquid height was equal to the tank diameter (T). Commonly, low-dispersed-phase systems and surfactants are used to reduce and eliminate coalescence in the system (Calabrese et al., 1986b; Wang and Calabrese, 1986; Sis and Chander, 2004; El-Hamouz, 2007). A surfactant solution was prepared with dilution of approximate 0.3%� (w/w) of sodium dodecyl sulfate (SDS) in the water before adding to the tank (El-Hamouz, 2009; Zainal Abidin et al., 2014). After

around 30 min of agitation, 1%�w/w of oil was added slowly into the mixture of water and surfactant (El-Hamouz, 2009). The Reynolds number greater than 104 in the mixing vessel results in turbulent flow regime (Kraume and Zehner, 2001). Therefore, impeller speeds of 5.0, 6.0, and 7.0 rps were chosen to create fully turbulent conditions within the vessel with Reynolds numbers between 88,445 and 123,823. These impeller speeds prevent surface aeration during drop size measurements (El-Hamouz, 2007). The mixing process was also examined at agitation speeds of 8.0 and 9.0 rps for CB180. Samples were withdrawn from the tank and analyzed. The sampling points were the same for all measurements, which were 0.02 m above the impeller level (O’Rourke and MacLoughlin, 2005; El-Hamouz, 2007). This sampling point was chosen as there was higher drop break-up rate around the impeller region compared to the other regions in the stirred vessel (Zainal Abidin et al., 2014). 0.002 m3 (200 ml) of sample were taken

Table II. Physical properties of the dispersed and continuous phases

Fluid Viscosity (kg/ms)

Density (kg.m−3)

Refractive index

Surface tension

(mN m−1) Water 0.001 998.00 1.3331 68.88 Cooking

palm oil 0.08198 890.00 1.4645 31.44

Table III. Type of impellers investigated in the experimental part

No. Impeller type Outer Dia.

(D) (m) Curvature

angle Central disk

size (m) Blade

length (m) Blade

thickness (m) D/T P0

1 6-curved blade with 1/2 D central disk (CB1/2D) 0.133 180 1/2D (0.067) 0.0053 0.002 1/3 3.9 2 6-curved blade with 1/4 D central disk (CB 1/4D) 0.133 180 1/4D (0.033) 0.0043 0.002 1/3 4.1 3 6-curved blade with 3/4 D central disk (CB3/4D)

(CB180°) 0.133 180 3/4D (0.1) 0.0033 0.002 1/3 3.4

4 6-curved blade with 3/4 D central disk (CB160°) 0.133 160 3/4D (0.1) 0.0033 0.002 1/3 3.7 5 6-curved blade with 3/4 D central disk (CB140°) 0.133 140 3/4D (0.1) 0.0033 0.002 1/3 4.9

Fig. 2. Schematic diagram of the experimental setup: A. load cell; B. connector; C. rod; d. distance from the motor to the central rod; E. metal sheet; F. motor; G. Weight (Afshar Ghotli et al., 2013).

Effect of Various Curved-Blade Impeller Geometries 5

after 1 h of mixing. Sampling was repeated at the same time interval to prevent time-induced error (O’Rourke and MacLoughlin, 2005). The literature suggested using more than 500 drops to obtain more reliable results (Buffo and Alopaeus, 2016). Samples were analyzed with Malvern Mastersizer, an instrument encompassing laser diffraction technique (O’Rourke and MacLoughlin, 2005). The tanks and sampling tubes were taken apart and washed with detergent and acetone after the experiments. Then, they were repeatedly washed with water to remove any trace contaminants.

Dynamic Light Scattering Laser Diffraction Several experimental methods can be employed to determine and monitor the mean drop size (d32) and drop size distributions (DSD) in mixing vessels (Pacek et al., 1994; Ribeiro et al., 2004; O’Rourke and MacLoughlin, 2005; Maaß et al., 2012; Abidin et al., 2013). Most of these methods have been applied for systems with low-volume dispersed phase (Pacek et al.,

1999). Based on Table I, several works using the laser diffrac-tion method have been reported. In this study, Malvern laser light scattering instrument (Malvern Instruments, Worcester-shire, UK) was used to measure the drop size distribution of water/oil dispersions. Drop size is defined based on the measurement and interpretation of angular distribution of dif-fracted light by drops using Mie’s theory (El-Hamouz, 2007). Malvern Mastersizer 2000 is capable of detecting particles and drops with sizes ranging from 0.02–2000 µm with an accuracy of �1%� on volume median diameter using a single-lenses system. The accuracy of the instrument was checked using deionized water (El-Hamouz et al., 2009).

The samples were diluted with deionized water before being vanalyzed by Malevern 2000. All the measurements were performed at 297 K. The analysis was then carried out by the software associated with Malvern 2000. This method has been applied successfully by many researchers (Boye et al., 1996). The Sauter mean diameter (d32) is defined based on

Fig. 3. The Schematic diagram of different types of impellers utilized in this work: (1) 6 CB 140° and 160° with 3/4 D central disk size; (2) 6 CB 180° with 3/4 D central disk size; (3) 6 CB 180° with 1/2 D central disk size; (4) 6 CB 180° with 1/4 D central disk size (Afshar Ghotli et al., 2013).

6 R. A. Ghotli et al.

Expression 5. The precision of the particle size measurements is �5%� (Skelland and Kanel, 1992; Zaldívar et al., 1996; Srivastava et al., 2000; Paul et al., 2004).

Statistical Analysis The Response Surface Methodology (RSM) is confirmed as a reliable mathematical and statistical technique for modeling, analysis, and determination of regression model equations and operational conditions based on quantitative data of appropriate experiments in some chemical engineering processes (Montgomery, 2001; Sahu et al., 2009; Vargas et al., 2010; Arami-Niya et al., 2011; Baroutian et al., 2011a; Baroutian et al., 2011b; Shafeeyan et al., 2012; Afshar Ghotli et al. 2013; Rahim Pouran et al., 2015; Ricardo et al., 2008). Therefore, in this study, RSM, using a Central Composite Design (CCD) was used for designing the experiments with a minimum number of laboratory experiments to analyze the

effects of curvature angle, impeller rotational speed and central disk size on d32. With this technique, a core factorial is created that forms a cube with sides that are two coded units in length (from −1 to þ1 as noted in Table IV). Table IV illustrates the coded values of each factor level consisting of agitation speed, X1, angle size, X2 and disc size, X3, with their coded levels for the CCD.

Results and Discussion

Effect of Impeller Speeds on the Mean Drop Size The drop size distribution curves at different impeller speeds for CB 180 were plotted in Figure 4. The curves show the typical evolution of drop sizes during dispersion. It was observed that higher agitation speed produced finer droplets, regardless of the size of central disk or curvature angle. The drop size distri-bution plots evidently illustrate that the peak in the distribution curves became narrower and shifted toward finer drop size with increasing impeller speed (Figure 4). Furthermore, the volume frequency of smaller drops increased at higher agitation speed. The results were in a good agreement with the reported data in the literature (Mlynek and Resnick, 1972; Ribeiro et al., 2004; Lovick et al., 2005; Jakobsen, 2008). Similar trend was also observed for the other impellers. It was attributed to the influence of higher energy dissipation rate on the drop breakage rate at higher agitation speed. Increased impeller speed causes higher power input which increases the shear and turbulent energy dissipation rates. Consequently, there is increased external stress on drops which leads to higher breakage or smaller coalescence rates (Giapos et al., 2005; Sis et al., 2005; Zainal Abidin et al., 2014). Therefore, the drops are much coarser at lower speeds than at higher speeds. Increasing the speed causes break-up and increases the number of droplets with smaller size.

Effect of Agitation Speed on d32 at Increasing the Angle and Central Disk Size The d32 data for each impeller based on the curvature angle and central disk size were plotted as a function of impeller agitation speed N in Figure 5(a) and (b). Based on Figure 5(a), it was observed that d32 values reduced with decreased curvature angle from 180° to 140°.

The results showed that there was about 9–18%�smaller drop size at higher impeller speeds for all the studied impellers. This could be related to the power consumption rate of impellers.

Fig. 4. Typical drop size distribution curves in different speeds for CB 180.

Fig. 5. Sauter mean diameter vs agitation speed based on different; (a) curvature angle and (b) central disk size.

Table IV. Independent variables and their coded levels for the CCD

Variables Code

Coded variable level

−1 0 1 Agitation speed (rps) X1 5 6 7 Angle size (Degree) X2 140 160 180 Disc size (m) X3 0.03 0.07 0.1

Effect of Various Curved-Blade Impeller Geometries 7

Higher power consumption resulted in smaller drop sizes. This behavior could also be described by circulation time and the impeller designs. Decrease in curvature angle from 180° to 140° resulted in wider blade and larger pumping volume. There-fore, the circulation time was shorter and the rate of breakage increased. Figure 5(b) also shows that smaller drop size was pro-duced by reducing the central disk size from 3/4 D to 1/4 D. A reduction in central disk size caused higher power consumption and smaller drop size. According to the literature, smaller cen-tral disk size creates more turbulence in the system and results in shorter circulation time and faster breakage in drops. In this study, different slopes were obtained for lines which were fitted to the datasets of different angles and central disk sizes. The slopes represent the changes in droplet size as a function of agi-tation speed. In the case of curvature angles, the values of −10.5 for all curvature angles showed the same trend for different studied curvature angles. Although the CB140 produced smaller droplet sizes, the same slopes showed nearly the same variation for each impeller at different agitation speeds. On the other hand, the values of −10.5, −11.0 and −13.5 confirmed that the largest slope was obtained for CB1/4D. Larger change was observed for smaller central disk size through increasing the agi-tation speed. The difference in the slope values could be explained by the effect of drop diameter on the rate of drop breakage. The stability of smaller drops is higher than that of the larger drops. Probability of drop breakage increased for larger drops due to larger drop break-up area (Zainal Abidin et al., 2014).

Development of the Regression Model The empirical relationships between the response (d32) and three independent variables: i) agitation speed; ii) curvature angle; and iii) central disk size were developed using Central Com-posite Design. Five trials were conducted at the center point to establish the experimental error. Tables V and VI present the experimental design and corresponding results. Y1 and Y2 show the responses (d32) based on the two considered cate-gories. The software suggested a quadratic model for both Y1

and Y2. Equation 7 presents the relationship between d32 (Y1) with agitation speed (X1) and curvature angle (X2) while Equation 8 shows the relationship between d32 (Y2) with agitation speed (X1) and central disk size (X3) in terms of coded factors. Positive terms indicate synergistic effects whereas nega-tive terms signify antagonistic effects (Ahmad et al., 2009; Houshmand et al., 2011). Both quadratic models were chosen as there were lower standard deviation and higher R-squared values. The R-squared value for Y1 and Y2 was 0.9750 and 0.9916 respectively. High R-squared values signify a good agreement between the predicted and experimental values (Houshmand et al., 2011). Furthermore, the close proximity of the points to the line, as shown in Figure 6, also proves the reliability of the developed models.

Y1 ¼ 152:55 � 9:83X1 þ 6:83X2 þ 3:57X 21

ð7Þ

Y2 ¼ 141:90 � 11:67X1 þ 15:17X3 þ 2:86X 21 þ 4:36X 2

3 ð8Þ

Statistical Analysis on Experimental Data Determination of coefficients and adequacy of the suggested models was assessed by Analysis of Variance (ANOVA). The statistical results and significance of the models for predicting droplet size of curved-blade impellers with different curvature angles and central disk sizes are given in Tables VII and VIII, respectively. The total variation was divided according to the sources of variation to examine the hypotheses based on the parameters of the model (Gonen and Aksu, 2008; Shafeeyan et al., 2012). Tables VII and VIII justify the significance both quadratic models with F-value of 54.55 and 164.32. Besides, it was found that Prob > F was less than 0.05 which implied that the terms were significant (Ahmad et al., 2009).

For the Y1 model, agitation speed (X1), angle (X2) and X12

were significant model terms while for the Y2 model, agitation speed (X1), central disk size (X3), X12 and X32 were significant model terms. In addition, a high value of Prob > F for X1.X2 (0.2142) and X22 (0.6213) showed that the influence of these terms on the response were insignificant. Adequate precision measures the signal to noise ratio and a ratio greater than 4 is desirable. The ratios of 26.807 and 46.964 for model equations

Table VI. Experimental design matrix and response results for various central disk sizes

Run Type Agitation speed,

X1 (rps) Central disk size, X3 (m)

d32, Y2 (µm)

1 Axial 7 0.0665 134 2 Axial 6 0.0997 161 3 Factorial 5 0.0333 147 4 Center 6 0.0665 140 5 Axial 6 0.0333 132 6 Factorial 5 0.0997 175 7 Factorial 7 0.0997 154 8 Center 6 0.0665 145 9 Axial 5 0.0665 156

10 Center 6 0.0665 142 11 Factorial 7 0.0333 120 12 Center 6 0.0665 142 13 Center 6 0.0665 140

Table V. Experimental design matrix and response results for various angles

Run Type Agitation speed,

X1 (rps) Angle, X2 (degree)

d32, Y1 (µm)

1 Center 6 160 153 2 Center 6 160 151 3 Factorial 7 180 154 4 Factorial 5 180 172 5 Axial 6 180 161 6 Factorial 5 140 162 7 Center 6 160 156 8 Center 6 160 152 9 Center 6 160 151

10 Axial 7 160 147 11 Axial 6 140 145 12 Axial 5 160 165 13 Factorial 7 140 139

8 R. A. Ghotli et al.

(8) and (9) indicated an adequate signal, which showed that these models were applicable to navigate the design space.

The Effect of Curvature Angles on Drop Size The effects of curvature angles and agitation speed on Sauter Mean Diameter (d32) for curved-blade impellers are presented in Figure 7. The most important observation of the current study was that Sauter Mean Diameter always increased with increase in blades curvature angle. A decrease from 180° to 140° may reduce the drop size as much as 9%, 10%, and 10%�at agitation speed of 5, 6 and 7 rps, respectively.

The drop size was found to be in the range of 175 to 139 µm for different curvature angles and speeds. An increase in drop size with increasing curvature angles was expected due to the smaller power number and accordingly smaller power consumption for the larger curvature angles. Therefore, the rate of breakage increased with decreasing curvature angle, leading to smaller drop size (Zhou and Kresta, 1996; Giapos et al., 2005; Rajapakse, 2007). However, the results showed that the Sauter Mean Diameter values were relatively close to each other at the various rotational speeds.

Fig. 6. Predicted vs. actual value of Drop size (d32) in terms of: (a) Curvature angle and (b) Central disk size.

Table VII. Analysis of variance of the linear model for drop size response (Curvature Angle)

Source Sum of squares Degree of freedom (dF) Mean of square F-Value p-Value (prob>F) Remarks Model 913.78 5 182.76 54.55 <0.0001 Significant X1 580.17 1 580.17 173.18 <0.0001 Significant X2 280.17 1 280.17 83.63 <0.0001 Significant X1X2 6.25 1 6.25 1.87 0.2142 Insignificant X12 35.18 1 35.18 10.50 0.0142 Significant X22 0.89 1 0.89 0.27 0.6213 Insignificant Residual 23.45 7 3.35 Lack of fit 6.25 3 2.08 0.48 0.7110 Insignificant Pure error 17.20 4 4.30 Cor total 937.23 12 R-Squared 0.975

Table VIII. Analysis of variance of the quadratic model for drop size response (Central disk size)

Source Sum of squares Degree of freedom (dF) Mean of square F-Value p-Value (Prob>F) Remarks Model 2324.50 5 464.90 164.32 <0.0001 Significant X1 816.67 1 816.67 288.65 <0.0001 Significant X3 1380.17 1 1380.17 487.82 <0.0001 Significant X1X3 9.00 1 9.00 3.18 0.1177 Insignificant X12 22.62 1 22.62 8.00 0.0255 Significant X32 52.55 1 52.55 18.57 0.0035 Significant Residual 19.80 7 2.83 Lack of fit 3.00 3 1.00 0.24 0.8657 Insignificant Pure error 16.88 4 4.20 Cor total 2344.31 12 R-Squared 0.992

Effect of Various Curved-Blade Impeller Geometries 9

The Effect of Central Disk Sizes on Drop Size Figure 8 illustrates the effect of central disk size and agitation speed on the Sauter Mean Diameter (d32) for curved-blade impellers. It was observed that the Sauter Mean Diameter increased with increase in central disk size. A decrease from

3/4 D to 1/4 D reduced the drop size as much as 16%, 18%�

and 22%� at agitation speed of 5, 6, and 7 rps, respectively. Based on the design matrix (Table VI), drop size was in the range of 175 to 120 µm for the studied central disk sizes and agitation speeds. An increase in drop size with central disk size

Fig. 7. The effect of various curvature angle and agitation speed on mean drop size (d32) for curved blade impellers.

Fig. 8. The effect of various central disk sizes and agitation speeds on mean drop size (d32) for curved blade impellers.

10 R. A. Ghotli et al.

was expected since bigger central disk size leads to smaller power numbers. The results indicate that the effect of central disk size on droplet sizes is much bigger than the curvature angle.

Effect of Average Energy Dissipation Rate on d32 at Increasing the Angle and Central Disk Size The d32 data for each impeller based on the curvature angle and central disk size were also plotted as a function of average energy dissipation rate (e) in Figure 9(a) and (b). Figure 9(a) illustrates that the d32 values decreases with increasing e, while the value of e was observed to have increased with decrease in curvature angle from 180° to 140°. Additionally, Figure 9(b) presents smaller drop size by reducing the central disk size from 3/4 D to 1/4 D. The power consumption and e for impellers with higher power number were higher at a fixed agitation speed and impeller diameter.

The CB180 had the lowest power number value compared to the other curved blades, agreeably, power number decreases with the enhancement in the curvature angle and smaller curva-ture angles have wider curvature and consequence in larger swept areas which leads to higher power consumption (Afshar Ghotli et al., 2013). On the other hand, higher power consump-tion of CB140 compared to the others resulted in smaller drop sizes.

In the case of different central disk size, the results indicated the curved blade impeller with bigger central disk size had the lowest power number. Decreasing the central disk size while the impeller diameter is constant results in higher net swept area and consequently higher power consumption. The flow instabil-ities can be declined with increase in central disk size (Vasconceloset al., 1999). Hence, the power consumption for the impellers with bigger central disk size were lower (Afshar Ghotli et al., 2013).

Conclusion A batch system consisting of 1%�oil in water was used to study the effects of curvature angles and central disk size of six curved-blade impellers on the Sauter Mean Diameter and drop size distribution. As expected, the study showed that higher agi-tation speed led to smaller drop size (d32). A decrease in curva-ture angle from 180° to 140° reduced d32 as much as 9%, 10%, and 10%�at 5, 6, and 7 rps, respectively. Besides, the decrease in

central disk size from 3/4 D to 1/4 D reduced the drop size up to 16%, 18%, and 22%� at agitation speed of 5, 6, and 7 rps, correspondingly. Two models were derived based on the experimental value through the DOE software and the adequacy of the models was statistically assessed with ANOVA. It can be concluded that at the same energy dissipation rate the impeller with smaller central disk size and curvature angles produces finer droplet sizes, in the hydrodynamic point of view. Moreover, the effect of variations in central disk sizes on droplet sizes is more significant than the curvature angles. The outcome of this research could be helpful in designing impellers for various mixing processes, especially liquid–liquid mixing.

Funding The authors are grateful for the financial support provided by the University of Malaya Research Grant (UMRG) (RP010C- 13SUS) and the Department of Chemical Engineering, University of Malaya for providing the necessary facilities to conduct the research.

Nomenclature T Tank Diameter (m) D Impeller Diameter (m) C Impeller Clearance (m) B Baffle width (m) We Webber Number (Dimensionless) N Agitation Speed (rps) qc Continuous Phase Density (kg.m−3) l Viscosity (kg/ms) r Interfacial tension (mN m−1) P Power Consumption (kg.m2 s−3) P0 Impeller Power number (Dimensionless) vT Liquid Volume in the vessel (m3) d32 Sauter mean diameter (µm) dmax Maximum drop size diameter (µm) ni Number of drops di Nominal drops diameter (µm)

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Effect of Various Curved-Blade Impeller Geometries 13