ISSN 1451 - 9372(Print) ISSN 2217 - 7434(Online) JANUARY ... No1_p1-150_Jan-M… · Djordjevic, Modelling of polyester fabric dyeing in the presence of ultrasonic waves ..... 131

ISSN 1451 - 9372(Print)ISSN 2217 - 7434(Online)JANUARY-MARCH 2017Vol.23, Number 1, 1-150

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CONTENTS

Marija Mihajlović, Marija Stanojević, Mirjana Stojanović, Je-lena Petrović, Jelena Milojković, Marija Petrović, Zo-rica Lopičić, To what extent do soft mechanical activation and process parameters increase the efficiency of different zeolite/phosphate rock fertilizer mixtures? ................................................................................ 1

Fatih Ilhan, Kaan Yetilmezsoy, Harun Akif Kabuk, Kubra Ulu-can, Tamer Coskun, Busra Akoglu, Evaluation of operational parameters and its relation on the stoichiometry of Fenton’s oxidation to textile was-tewater .................................................................................. 11

Javad Ahmadishoar, S. Hajir Bahrami, Barahman Movas-sagh, Seyed Hosein Amirshahi, Mokhtar Arami, Rem-oval of Disperse Blue 56 and Disperse Red 135 dyes fromaqueous dispersions by modified montmorillonite nanoclay ............................................................................... 21

Alireza Ebrahiminezhad, Yahya Barzegar, Younes Ghasemi, Aydin Berenjian, Green synthesis and characterization of silver nanoparticles using Alcea rosea flower extract as a new generation of antimicrobials .................................. 31

Aleksandra Mišan, Bojana Šarić, Ivan Milovanović, Pavle Jovanov, Ivana Sedej, Vanja Tadić, Anamarija Mandić, Marijana Sakač, Phenolic profile and antioxidant properties of dried buckwheat leaf and flower extracts ........ 37

Yajing Zhang, Yu Zhang, Fu Ding, Kangjun Wang, Xiaolei Wang, Baojin Ren, Jing Wu, Synthesis of DME by CO2 hydrogenation over La2O3-modified CuO–ZnO– –ZrO2/HZSM-5 catalysts ....................................................... 49

Tatjana Kaluđerović Radoičić, Nevenka Bošković-Vragolović, Radmila Garić-Grulović, Mihal Đuriš, Željko Grbavčić, Friction factor for water flow through packed beds of spherical and non-spherical particles ................................... 57

Hai-Peng Gou, Guo-Hua Zhang, Kuo-Chih Chou, Prepar-ation of titanium carbide powder from ilmenite con-centrate ................................................................................. 67

Milana M. Zarić, Mirko Stijepovic, Patrick Linke, Jasna Stajić-Trošić, Branko Bugarski, Mirjana Kijevčanin, Targeting heat recovery and reuse in industrial zone ............................ 73

Sandra Raquel Kunst, Lilian Vanessa Rossa Beltrami, Mari-elen Longhy, Henrique Ribeiro Piaggio Cardoso, Tiago Lemos Menezes, Célia de Fraga Malfatti, Effect of diisodecyl adipate concentration in hybrid films applied to tinplate .............................................................................. 83

CONTENTS Continued Milovan Janković, Snežana Sinadinović-Fišer, Olga Goveda-rica, Jelena Pavličević, Jaroslava Budinski-Simendić, Kinetics of soybean oil epoxidation with peracetic acid formed in situ in the presence of an ion exchange resin: Pseudo-homogeneous model .................................... 97

Salah H. Aljbour, Sultan A. Tarawneh, Adnan M. Al-Harah-sheh, Evaluation of the use of steelmaking slag as an aggregate in concrete mix: A factorial design approach .... 113

Dušan Lj. Petković, Miloš J. Madić, Goran M. Radenković, The effects of passivation parameters on pitting potential of biomedical stainless steel ................................ 121

Marija Kodric, Sandra Stojanovic, Branka Markovic, Dragan Djordjevic, Modelling of polyester fabric dyeing in the presence of ultrasonic waves ............................................. 131

Aishi Zhu, Shanshan Liu, Kanfeng Wu, Chuan Ren, Maoqian Xu, Comparing of hot water and acid extraction of polysaccharides from proso millet ...................................... 141

Activities of the Association of Chemical Engineers of Serbia are supported by: - Ministry of Education, Science and Technological Development, Republic of Serbia - Hemofarm Koncern AD, Vršac, Serbia - Faculty of Technology and Metallurgy, University of Belgrade, Belgrade, Serbia - Faculty of Technology, University of Novi Sad, Novi Sad, Serbia - Faculty of Technology, University of Niš, Leskovac, Serbia - Institute of Chemistry, Technology and Metallurgy, University of Belgrade, Belgrade, Serbia

Chemical Industry & Chemical Engineering Quarterly

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Chem. Ind. Chem. Eng. Q. 23 (1) 19 (2017) CI&CEQ

1

MARIJA MIHAJLOVIĆ

MARIJA STANOJEVIĆ

MIRJANA STOJANOVIĆ

JELENA PETROVIĆ

JELENA MILOJKOVIĆ

MARIJA PETROVIĆ

ZORICA LOPIČIĆ

Institute for Technology of Nuclear

and Other Mineral Raw Materials,

Belgrade, Serbia

SCIENTIFIC PAPER

UDC 631.82/.85:549.67:553:66

https://doi.org/10.2298/CICEQ150622047M

TO WHAT EXTENT DO SOFT MECHANICAL ACTIVATION AND PROCESS PARAMETERS INCREASE THE EFFICIENCY OF DIFFERENT ZEOLITE/PHOSPHATE ROCK FERTILIZER MIXTURES?

Article Highlights

• Effect of mechanical activation on different exchange-fertilizer mixtures was investigated

• Substantially increased P content than that of the similar non-activated mixtures was

observed

• Influence of different process parameters was explored using multivariate statistics

• The results confirmed a high fertilization potential of the selected mechanically acti-

vated substrates

Abstract

In order to obtain effective mineral fertilizer, different mixtures of phosphate rock

(PR) with natural clinoptilolite (Cp) and NH4+ saturated clinoptilolite (NH4-Cp)

were subjected to soft mechanical activation. Mean concentrations of P released

from mechanically activated (MA) substrates, MACp/PR and MANH4-Cp/PR,

ranged between 2.81-3.19 mg L-1 and 2.02-7.74 mg L-1, respectively. These are

10 to 15 times higher P concentrations than those released from the corres-

ponding non-activated mixtures. Solution Ca2+, K+, Na+ and Mg2+ concentrations

varied according to the composition of the mixtures and the contact time

between the two minerals within their optimal values required for plant growth.

The obtained results suggest that the efficiency of the NH4-Cp/PR mixtures can

be significantly increased by the proposed mechanical activation. Influence of

process parameters on the observed concentrations of nutrients was shown

using multivariate statistics. The highest fertilization potential demonstrated

MANH4-Cp/PR mixture with the largest NH4-Cp share and the longest proposed

mixing time.

Keywords: exchange-fertilizer mixtures, phosphate rock, clinoptilolite, mechanical activation, multivariate analysis.

Correct use of fertilizers and natural resources

contributes to environmental sustainability. Inorganic

fertilizers account for about 80% of all phosphate

applications, of which more than 99% originate from

the phosphate rock (PR) [1]. Traditional technologies

for production of high soluble mineral fertilizers from

natural PR include acid treatments causing environ-

mental contamination and eutrophication [2]. To avoid

Correspondence: M. Mihajlovic, Central Laboratory for Testing,

Institute for Technology of Nuclear and Other Mineral Raw

Materials, Franše Deperea 86, 11000 Belgrade, Serbia. E-mail: [email protected] Paper received: 22 June, 2015 Paper revised: 12 November, 2015 Paper accepted: 7 December, 2015

these issues, it is necessary to find new methods to

obtain soluble fertilizers with less negative impact on

the environment. The direct application of PR seems

to be one of the effective strategies with lower energy

usage and production costs. However, the low sol-

ubility, therefore the low attainability of nutrients for

the plant growth, is the main disadvantage of the PR

direct utilization [3,4]. One of the proposed approaches,

in aim to increase dissolution of PR, is the application

of combined zeolite/PR mixtures. The addition of nat-

ural zeolites to the PR, usually clinoptilolite (Cp), imp-

roves the PRs agrochemical effect [5,6]. The high

cation-exchange capacity, soil texture enrichment and

water-retention ability are main features of zeolites,

M. MIHAJLOVIĆ et al.: TO WHAT EXTENT DO SOFT MECHANICAL ACTIVATION… Chem. Ind. Chem. Eng. Q. 23 (1) 19 (2017)

2

which determine their successful agricultural appli-

cations [7-11]. Characteristics of fertilizer mixtures

could be further enhanced through modification of

natural zeolites with nutrient elements such as NH4+

[12-14]. Thus, the zeolites, beside the increase of the

PR solubility can become a source of N, Ca2+ and K+

required for plant growth [6,14].

Another approach for increasing PR solubility is

its mechanical activation that produces physical and

chemical changes in close surface regions where

mechanical energy leads to contact between the

solids [15,16]. The extensive knowledge of solid-state

mechanochemistry has been expanded with the obs-

ervations of the influence of high pressure and shear

on the rate of chemical solid-state reactions [16].

However, some authors note that the particle size

reduction due to mechanical activation affects the

PRs reactivity to a greater extent than the resulting

deformations of its crystal structure [17]. Subse-

quently, it has been found that the efficiency of

exchange-fertilizer mixtures may be further enhanced

by their mechanical activation [18-22]. To what extent

this procedure increases the efficiency of different

zeolite/PR mixtures, so that its inclusion into the pro-

duction process could be considered rational, was the

subject of our survey. Therefore, the influence of soft

mechanical activation on improvement of the PR dis-

solution in the presence of two types of clinoptilolites,

natural (Cp) and partially NH4+ saturated (NH4-Cp)

was studied. The changes in mineralogical structure,

and concentrations of plant available nutrients rel-

eased from both MACp/PR and MANH4-Cp/PR mix-

tures, were investigated and compared with that of

the corresponding non-activated mixtures. Differ-

ences in concentrations of the released nutrients and

their dependencies on the process parameters

between various MA-substrates were analyzed using

multivariate statistical methods.

EXPERIMENTAL

Characterization of materials

Natural phosphate ore, PR, from deposit Lisina,

near Bosilegrad in Serbia with average content of 9%

P2O5 and zeolite tuff (with >75% of clinoptilolite (Cp))

from deposit Baia Mare, Romania, were used for the

preparation of the mixtures. Selected characteristics

of the Cp and PR samples are presented elsewhere

[6]. To obtain the NH4-exchanged Cp, a partial sat-

uration of the Cp with NH4+ at 1:7.5 ratios was per-

formed, according to the procedure described by

Mihajlović et al. [23].

Powder X-ray diffraction (XRD) was used to

determine the phase composition of the Cp/PR and

NH4−Cp/PR mixtures before and after mechanical

activation. The XRD patterns were obtained on a

Philips PW-1710 automated diffractometer using a Cu

tube operated at 40 kV and 30 mA. The instrument

was equipped with a diffracted beam curved graphite

monochromator and a Xe-filled proportional counter.

The diffraction data were collected in the 2 Bragg

angle range from 5 to 60°, counting for 2 s (qualitative

identification) at every 0.02° step. The divergence and

receiving slits were fixed at 1 and 0.1 units, res-

pectively. The XRD measurements were performed at

room temperature in a stationary sample holder.

Experimental procedure

Both groups of mixtures (Cp/PR and NH4-Cp/

/PR), in three replicates, were assembled in three

ratios of Cp and the PR; 5:1 10:1 and 15:1. Each

mixture originally contained 4 g of PR and the corres-

ponding share of zeolite (20, 40 and 60 g, respect-

ively). The mechanical activation of the mixtures was

carried out in a vibrating ring-mill (KHD, Humboldt

Wedag, AG). To avoid sticking, the period of mech-

anical activation was of 30 s per sample at room tem-

perature. After mechanical activation, mixtures were

placed in 300 ml volumetric flasks. Then, in each vol-

umetric flask was added 200 ml of distilled water and

mixtures were shaken on a rotary shaker for 24, 48

and 72 h at 220 rpm. In the resulting solutions, after

draining, the concentrations of Ca2+, K+, Na+ and Mg2+

were determined using a Perkin Elmer AAS 703

atomic absorption spectrometer. The concentration of

P from the solution was determined by colorimetry

[24] using a Spekol 1300 UV–Vis spectrophotometer.

Exploratory data analysis

Descriptive statistical analyses of the results

were expressed as the mean ± standard deviation

(SD). Analysis of variance (ANOVA) and the following

posthoc Tukey’s HSD were performed to determine

whether there are any significant differences between

the average values of the released nutrients between

the groups.

Multivariate statistical analysis

Principal component analysis (PCA) as a multi-

variate analytical method was used to display the

data in a multidimensional space, where the variables

determine the axes [25]. These axes are projected

into a few principal components (PCs), which are

linear combinations of the original variables and

define the maximum variation within the data. The first

principal component (PC1) accounts for the largest


3

variance in the data set. The following principal com-

ponents account for the maximum of the remaining

variance in the data set. Each PC is characterized by

2-D scores plot maps, which shows a data overview

and similarities or dissimilarities within the data. In

this study PCA was employed to establish: i) the

differences between nutrients release from the pure

PR and mechanically activated substrates (MACp/PR

and MANH4-Cp/PR mixtures) by calculating score

plots and ii) the influence of process parameters (zeo-

lite/PR ratio and the mixing time) on the content of

nutrients (P, Ca2+, K+, Na+ and Mg2+) in the solutions

of both MA-mixtures by calculating correlation loading

plots. All PCA calculations were carried out by PLS

ToolBox, (Eigenvectors Inc., v. 7.9), for Matlab 7.12.0

(R2011a) and The Unscrambler (version 9.7, CAMO

Process AS, Oslo, Norway). PCA was applied by

using a 0.95 confidence level for Q and T2 Hotelling

limits for outliers and a singular value decomposition

algorithm (SVD). All data were auto scaled prior to

any PCA analysis.

RESULTS AND DISCUSSION

Mineral characterization

The mineralogical compositions of Cp/PR and

NH4-Cp/PR substrates (with highest zeolite share,

15:1) before and after mechanical treatment are

shown in Figure 1. All mixtures irrespective of the

composition were of similar mineralogical structure

(for which XRD patterns for mixtures at 5:1 and 10:1

Cp/PR ratio are omitted) and contained Heu-type

zeolite, quartz, plagioclase, muscovite, and apatite.

The crystallinity degree of minerals between non-act-

ivated and MA samples remained unchanged. Also,

systematic shifting of diffraction maximums of the

dominant mineral in the mixture, Cp, was rather neg-

ligible. Crystallite size for Heu-type zeolite measured

at (020) and (200) diffraction maximums were: Cp/PR

for <D> = 212 Å, MACp/PR for <D> = 283 Å, NH4-

-Cp/PR for <D> = 207 Å and MANH4-Cp/PR for <D> =

= 227 Å. The observed differences were due to the

presence of larger quantities of quartz and plagio-

clases in the Cp/PR and NH4Cp/PR, in relation to the

corresponding MA-mixtures, respectively (Figure 1a

and b).

The influence of mechanical activation on nutrient

release from the Cp/PR fertilizer mixtures

Solution P and Ca2+ concentrations released

from the MA-mixtures were compared with the con-

centration of the same elements leached from corres-

ponding non-activated mixtures [6], subject to their

composition and mixing time (Figure 2).

Figure 1. Diffractograms of Cp/PR and MACp/PR mixtures (a) and NH4-Cp/PR and MANH4-Cp/PR mixtures (b).


4

The concentrations of P released from the

MACp/PR mixtures ranged between 2.81 and 3.19

mg L-1 (Figure 2a). This is on average fifteen times

higher P concentration than those released from the

corresponding non-activated mixtures (0.15 to 0.26

mg L-1) [6]. The observed increase in P concentra-

tions due to soft mechanical activation of the mixtures

was in accordance with the expected upward trend for

P found in the literature [21]. Solution Ca2+ concentra-

tions gradually increased with the increase of the

MACp/PR ratio and the mixing time and were

between 11 and 22.5 mg L−1. However, Cp/PR and

MACp/PR substrates had similar solution Ca2+ con-

centration implying that mechanical activation did not

affect the changes in solution Ca (Figure 2b).

Solution K+, Na+ and Mg2+ concentrations rel-

eased from the MACp/PR and the comparable Cp/PR

mixtures [6] are shown in Figure 3. The increase in

MACp/PR ratio and contact time caused a slight inc-

rease of K+, Na+ and Mg2+ concentrations in the sol-

ution. Solution K+ and Na+ concentrations of the

MACp/PR mixtures were between 5.66-7.99 mg L-1

and 11.7-17.15 mg L-1, respectively. This is on aver-

age 30% lower than that of the corresponding non-

activated mixtures. Petkova et al. [20] previously

reported that tribochemical activation of the Cp/PR

leads to decrease of the Cp ion-exchange ability

whereas at the same time dissolution of the PR has

been increased [20]. Solution Mg2+ concentration of

the activated mixtures ranged between 0.83-1.65 mg

Figure 2. Mean solution P (a) and Ca2+ (b) concentrations ± SD of the Cp/PR [6] and MACp/PR mixtures.

Figure 3. Mean solution K+, Na+ and Mg2+ concentrations ± SD of the Cp/PR [6] and MACp/PR mixtures at a) 5:1, b) 10:1 and

c) 15:1 zeolite/PR ratio.


5

L-1 or around twofold higher in relation to the corres-

ponding Cp/PR mixtures (Figure 3).

The influence of mechanical activation on nutrient

release from the NH4-Cp/PR fertilizer mixtures

Solution P and Ca2+ concentrations of the

MANH4-Cp/PR mixtures were compared with the con-

tent of same nutrients leached from corresponding

non-activated mixtures [6], subject to their compo-

sition and the mixing time (Figure 4).

The mean concentrations of P released from the

MANH4-Cp/PR mixtures were between 2.02 and 7.74

mg L-1, which is ten times higher than those released

from the corresponding non-activated mixtures (0.36-

0.82 mg L-1) [6]. Similarly, 10-15 times higher extract-

ion of P to the acid media from the MA-phosphorite

has been reported previously [21]. Figure 4a shows

that in addition to mechanical activation, an increase

of the content of partially saturated NH4–zeolite in the

mixture and duration of contact between the two min-

erals significantly contributes the solubility of the PR.

Solution Ca2+ concentrations released from the

MANH4-Cp/PR mixtures varied between 445 and

510.5 mg L-1. The differences between dissolved Ca

from the NH4-Cp/PR and MANH4-Cp/PR were most

evident at 5:1 zeolite/PR ratio (Figure 4b). At higher

zeolite/PR ratios, the content of Ca in the solutions

was in favor of the non-activated mixtures. Further-

more, the Ca2+ concentration slightly decreased with

the increase of mixing time and Cp/PR ratio, similar to

the non-activated mixtures previously reported by

Mihajlovic et al. [6]. These results support the con-

clusions from the literature that the released Ca2+

from the PR, compelled by cation exchange with

NH4+, was partially withdrawn by the zeolite before

equilibrium [20,26].

Solution K+ and Na+ concentrations of the

MANH4-Cp/PR mixtures ranged from 40 to 59.4 mg L-1

and from 7.14 to15.45 mg L-1, respectively (Figure 5).

The content of both elements in the solution slightly

increased with the Cp/PR ratio increase, but ranged

around similar values in relation to the mixing time.

Nevertheless, in comparison to the non-activated mix-

tures, the K+ and Na+ concentrations in MA-fertilizers

with saturated zeolite were increased by 30 and 50%,

respectively. This supports the utilization of soft

mechanical activation, aiming to increase the leach-

ing of K from the tested fertilizer mixtures. Solution

Mg2+ concentrations in the MANH4-Cp/PR substrates

proportionally increased with increasing of zeolite

share in the mixtures and the mixing time, and ranged

between 20 and 54.5 mg L-1, very much alike to that

of the corresponding non-activated mixtures [6].

Figure 4. Mean solution P (a) and Ca2+ (b) concentrations ± SD of the NH4-Cp/PR [6] and MANH4-Cp/PR mixtures.


6

ANOVA

One-way ANOVA was applied to compare the

means between the concentrations of nutrients rel-

eased from the MACp/PR and the MANH4-Cp/PR

mixtures of different composition at various mixing

times (Tables 1 and 2). Statistical differences were

determined at the p < 0.05 level, 95% confidence

limit, according to Tukey’s HSD test.

Calculated one-way Fisher’s statistic critical

value was Fcrit = 5.14 and Tukey’s HSD post hoc

critical value was 4.34 for each element in both group

of mixtures. These values numerically define the

levels of significant differences in nutrient content,

released from the various mixtures according to their

composition and mixing time. If the obtained statistical

value, F and/or Tukey HSD value (between labeled

groups) for each element is larger than its critical

value, the differences in concentrations of the rel-

eased nutrient from the compared groups of mixtures

are significant at 0.05 levels.

Conversely to solution P concentrations, the

differences between solution Ca2+, K+, Mg2+ and Na+

concentrations released from the MACp/PR sub-

strates, were found to be significant in relation to the

composition of the mixtures (Table 1). Regarding the

calculated Tukey HSD values from Table 1, the con-

centrations of K+ released from the MACp/PR mixture

at 5:1 ratio and Mg2+ released from the MACp/PR

mixture at 15:1 ratio, were found to be significantly

different in comparison to the other two groups. The

effect of mixing time did not affect the occurrence of

significant differences between solution concen-

trations of the elements within the groups (Table 1).

Significant differences in relation to the mixture

composition were also found for K+ and Mg2+ released

from MA-substrates with saturated Cp, while the con-

centrations of Na+ released from the mixtures with

lowest NH4+-Cp share were significantly different in

comparison to the other two groups. Differences in

composition of the mixtures did not significantly affect

solution Ca2+ and P concentration, while the effect of

mixing time in the MANH4-Cp/PR mixtures was sig-

nificant at the p < 0.05 level only for P (Table 2).

Figure 5. Mean solution K+, Na+ and Mg2+ concentrations ± SD of the NH4-Cp/PR [6] and MANH4-Cp/PR mixtures at a) 5:1, b) 10:1 and

c) 15:1 zeolite/PR ratio.

Table 1. ANOVA of the MACp/PR mixtures; F – one way Fisher`s statistic test; p-value - function of the observed sample results;

HSD – Tukey`s post hoc test

Parameter P ratio Ca ratio K ratio Na ratio Mg ratio Parameter P time Ca time K time Na time Mg time

F 2.07 85.75 76.14 84.97 8.57 F 1.94 0.09 0.06 0.09 0.22

p-Value 0.21 3.8610-5 5.4510-5 3.9710-5 0.02 p-Value 0.22 0.91 0.94 0.92 0.81

HSD|15:1-10:1| - 6.87 3.25 8.55 4.45 HSD|72h-48h| - - - - -

HSD|15:1-5:1| - 18.33 16.47 18.42 5.53 HSD|72h-24h| - - - - -

HSD|10:1-5:1| - 11.46 13.23 9.87 1.09 HSD|48h-24h| - - - - -

Table 2. ANOVA of the MANH4-Cp/PR mixtures; F – one way Fisher`s statistic test; p-value - function of the observed sample results;

HSD – Tukey`s post hoc test

Parameter P ratio Ca ratio K ratio Na ratio Mg ratio Parameter P time Ca time K time Na time Mg time

F 0.88 2.03 222.12 99.67 83.27 F 7.53 1.82 0.02 0.02 0.06

p-Value 0.46 0.21 2.3710-6 2.4910-5 4.210-5 p-Value 0.02 0.24 0.98 0.98 0.94

HSD|15:1-10:1| - - 11.904 0.32 8.72 HSD|72h-48h| 1.64 - - - -

HSD|15:1-5:1| - - 29.62 17.45 18.25 HSD|72h-24h| 5.36 - - - -

HSD|10:1-5:1| - - 17.71 17.13 9.53 HSD|48h-24h| 3.72 - - - -


7

Principal component analysis

PCA was performed in order to explore the main

variation patterns between the data for pure PR,

MACp/PR and MANH4-Cp/PR mixtures in relation to

the different process parameters (Figure 6). The PCA

score plot with samples labeled in accordance to the

zeolite/PR ratio is presented in Figure 6a and the

PCA score plot with samples labeled in accordance to

mixing time is presented in Figure 6b. PCA resulted in

a two-component model, which explains 93.27% of

total variance. The first principal component, PC 1,

accounted for 70.67% and the second one, PC 2, for

22.60% of the overall data variance. The addition of

more PCs did not change the classification of the

samples. Taking into account PC1 and PC2 score

values, pure PR samples are grouped in the lower-left

part of the PCA score plot, MACp/PR samples in the

upper-left and MANH4-Cp/PR are separated in the

right part of the PCA score plots.

The zeolite share in the fertilizer mixtures caused

greater clustering of different MACp/PR samples, vis-

ible along positive PC2 axis. However, the overall imp-

act of the substrate composition is more evident in the

case of the MANH4-Cp/PR mixtures with obvious greater

separation of the samples along the first component.

Also, the highest zeolite share has the strongest

influence on solution cation concentrations (Figure 6a).

The different mixing times better affected the

separation of the MANH4-Cp/PR mixtures than of the

MACp/PR and the PR alone which is visible along

PC1 axis (Figure 6b). Furthermore, the influence of

mixing time increases as the zeolite/PR ratio inc-

reases, implying that the strongest effect of mixing

time is observed for the substrates with 15:1 zeo-

lite/PR ratio. Contrary to this, the influence of the

mixing time has a very weak effect on the MACp/PR

substrates. This indicates that changes in process

parameters had a greater impact on the release of

nutrients from the MANH4-Cp/PR mixtures.

To obtain a better overview of the effects of

zeolite/PR ratio in the mixtures and mixing time on

solution cation concentrations, two additional principal

component analyses were performed on the subsets

of data: for MACp/PR and for MANH4−Cp/PR samples

using correlation loading plots (Figure 6c and 6d).

From the correlation loading plot calculated for

MACp/PR substrates, it can be observed that the

highest zeolite/PR ratio exhibits the strongest influ-

ence on solution P, Ca2+, K+, Na+, and Mg2+ concen-

trations. The P content in the solution correlates the

best with MACp/PR ratio of 15:1.

Figure 4. Mean solution P (a) and Ca2+ (b) concentrations ± SD of the NH4-Cp/PR [6] and MANH4-Cp/PR mixtures.


8

The correlation loading plot calculated for

MANH4–Cp/PR substrates shows that the highest

zeolite/PR ratio exhibits the strongest influence on

solution P, K+, Na+ and Mg2+ concentrations, while the

Ca content was in best correlation with the lowest

zeolite share in the mixture and the shortest mixing

time. Also, Ca is highly negatively correlated with P.

The best correlation between solutions P concentra-

tion, was found for 72 h mixing time and MANH4-

–Cp/PR ratio of 15:1.

CONCLUSION

The obtained results showed that the use of soft

mechanical activation of tested fertilizer mixtures fol-

lowed by inducted particle size reduction, although

did not produce significant structural changes in min-

erals, notably intensified the passage of nutrients to

the liquid media, especially P. Solution P concentra-

tions of both MA substrates were up to fifteen times

higher than that of the corresponding non-activated

mixtures, while the release of K+ from the MANH4-

–Cp/PR mixtures was increased by a third. Further-

more, all concentrations of the nutrients released from

the MANH4-Cp/PR substrates were within their opti-

mal values necessary for plant growth [27,28]. Sol-

ution Ca2+ concentrations were between 445 and

510.5 mg L−1, which suggests that the use of the

MANH4-Cp/PR mixtures may resolve a potential defi-

ciency of the Ca2+ in the solution, very common for

substrates of similar composition [6].

PCA revealed that the highest NH4-zeolite share

in the mixture had the most positive impact on the PR

dissolution. A growth of fertilization potential with

time, particularly of the MANH4-Cp/PR mixtures, was

observed. This supports the use of the selected

MANH4–Cp/PR mixture as a slow-release fertilizer,

very favorable for plants since fertilization can be per-

formed less frequently, which, besides the efficiency,

increases the cost-effectiveness of its utilization.

Acknowledgment

The authors are grateful to the Serbian Ministry

of Education, Science and Technological Develop-

ment for the financial support of this investigation

included in the project TR 31003, projects cycle

2011−2015.

REFERENCES

[1] L. Maene, Phosphate fertilizer production, consumption

and trade, the present situation and outlook to 2010, The

Sulphur Institute's 17th Sulphur Phosphate Symposium,

January 17-19, Boca Raton, FL, 1999

[2] M. Hart, B., Quin; M. Nguyen, J. Environ. Qual. 33 (2004)

1954−1972

[3] L. Leon, W. Fenster, L. Hammond, Soil Sci. Soc. Am. J.

50 (1986) 798−802

[4] S. Chien, In: Direct application of phosphate rock and

related technology:latest development and practical exp-

eriences; S. Rajan, S. Chien (Eds.), Special Publications

IFDCSP- 37, IFDC, Muscle Shoals, AL, 2003, pp. 50−62

[5] W. Pickering, N. Menzies, M. Hunter, Sci. Hortic. 94

(2002) 333−343

[6] M. Mihajlović, N. Perišić, L. Pezo, M. Stojanović, J.

Milojković, Z. Lopičić, M. Petrović, J. Agric. Food Chem.

62 (2014) 9965−9973

[7] D. Ming, E. Mumpton, in Minerals in Soil Environments,

2nd ed., J. Dixon, S. Weed (Eds.), Soil Science Society of

America, Madison, WI, 1989, pp. 873−911

[8] C. Cobzaru , S. Oprea, Chem. Ind. Chem. Eng. Q. 11

(2005) 206-212

[9] C. Cobzaru, C. Cibotaru, A. Rotariu, A. Marinoiu, S.

Oprea, Chem. Ind. Chem. Eng. Q. 15 (2009) 63−67

[10] M. Onyango, J. Kittinya, N. Hadebe, V. Ojijo, A. Ochieng,

Chem. Ind. Chem. Eng. Q. 17 (2011) 385−395

[11] M. Mihajlović, S. Lazarević, I. Janković-Častvan, B. Jokić,

Dj. Janaćković, R. Petrović, Chem. Ind. Chem. Eng. Q.

20 (2014) 283-293

[12] K. Barbarick, T. Lai, D. Eberl, Soil Sci. Soc. Am. J. 54

(1990) 911−916

[13] E. Allen, L. Hossner, D. Ming, D. Henninger,. Soil Sci.

Soc. Am. J. 57(1993)1368−1374.

[14] [14] E. Allen, L. Hossner, D. Ming, D. Henninger, Soil Sci.

Soc. Am. J. 60 (1996) 1467−1472

[15] S. Ibrahim, A. El-Midany, T. Boulos, Physicochem. Probl.

Miner. Process. 44 (2010) 63-78

[16] B. Fotoohi, A study of mechanochemical activation in

solid-state synthesis of advanced ceramic composites, A

thesis submitted to The University of Birmingham, 2010

[17] N. Shulga, Russ. J. Appl. Chem. 85 (2012)1297−1306

[18] L. Shumskaya, E. Kirillova, T. Yusupov, J. Mining Sci. 35

(1999) 96-100

[19] N. Petrova, V. Petkova, Bulgarian Chem. Commun. 43

(2011) 301–307

[20] V. Petkova, E. Serafimova, N. Petrova, Y. Pelovski, J

Therm. Anal. Calorim. 105 (2011) 535–544

[21] T. Yusupov, L. Shumskaya, J. Mining Sci. 38 (2002) 177–181

[22] T. Yusupov, L. Shumskaya, E. Kirillova, V. Boldyrev, J.

Mining Sci. 42 (2006) 189-194

[23] M. Mihajlović, N. Perisic, L. Pezo, M. Stojanovic, J.

Milojković, M. Petrović, J. Petrović, Clay Minerals 49

(2014) 735-745

[24] R. Koenig, C. Johnson, Ind. Eng. 14 (1942) 155−156

[25] K. Pearson, Philos. Mag. 2 (1901) 559−572

[26] T. Lai, D. Eberl, Zeolites 6 (1986) 129−132

[27] S. Tisdale, W. Nelson, J. Beaton, In: Soil fertility and

fertilizers; Macmillan Publ. Co., New York, 1985

[28] G. Bryson, A. Barker, Commun. Soil Sci. Plant Anal. 33

(2002) 759−777.


9

MARIJA MIHAJLOVIĆ

MARIJA STANOJEVIĆ

MIRJANA STOJANOVIĆ

JELENA PETROVIĆ

JELENA MILOJKOVIĆ

MARIJA PETROVIĆ

ZORICA LOPIČIĆ

Institut za tehnologiju nuklearnih i

drugih mineralnih sirovina, Franše

D'Eperea 86, 11000 Beograd, Srbija

NAUČNI RAD

U KOJOJ MERI BLAGA MEHANIČKA AKTIVACIJA I PROCESNI PARAMETRI POVEĆAVAJU EFIKASNOST RAZLIČITIH ZEOLIT/RUDA FOSFORA SMEŠA PRIRODNIH ĐUBRIVA?

U cilju dobijanja efikasnog mineralnog đubriva, različite smeše rude fosfora (PR) sa pri-

rodnim klinoptilolitom (Cp) i NH4+-klinoptilolitom (NH4-Cp) podvrgnute su blagoj meha-

ničkoj aktivaciji. Srednje koncentracije P otpuštene iz mehanički aktiviranih (MA) smeša,

MACp /PR i MANH4-Cp/PR, varirale su u rasponu od 2,81-3,19 mg L-1 i 2,02-7,74 mg L-1,

redom. Ovo su 10 do 15 puta vece koncentracije P od onih otpuštenih iz odgovarajucih

neaktiviranih smeša. Koncentracije Ca2+, K+, Na+ i Mg2+ u rastvorima smeša varirale su u

zavisnosti od sastava smeše i vremena kontakta dva minerala, a u okviru njihovih opti-

malnih vrednosti potrebnih za rast i razvoj biljaka. Dobijeni rezultati pokazuju da se efikas-

nost NH4-Cp/PR smeše može značajno povecati predloženim postupkom mehaničke

aktivacije. Uticaj procesnih parametara na sadržaj posmatranih nutrijenata prikazan je upo-

trebom multivarijantne statističke analize. Najveci potencijal đubrenja pokazala je MANH4-

-Cp/PR smeša sa najvecim udelom zeolita i najdužim predloženim vremenom mešanja.

Ključne reči: smeše đubriva, ruda fosfora, klinoptilolit, mehanička aktivacija, multi-

varijantna analiza.






11

FATIH ILHAN

KAAN YETILMEZSOY

HARUN AKIF KABUK*

KUBRA ULUCAN

TAMER COSKUN*

BUSRA AKOGLU

Department of Environmental

Engineering, Faculty of Civil

Engineering, Yildiz Technical

University, Davutpasa, Esenler,

Istanbul, Turkey

SCIENTIFIC PAPER

UDC 66.094.3:628.3:677

https://doi.org/10.2298/CICEQ150907048I

EVALUATION OF OPERATIONAL PARAMETERS AND ITS RELATION ON THE STOICHIOMETRY OF FENTON’S OXIDATION TO TEXTILE WASTEWATER

Article Highlights

• Operation conditions for implementation of the Fenton process were re-evaluated

• COD/H2O2, COD/Fe2+ and H2O2/Fe2+ ratios were optimized by physico-chemical studies

• The effects of important parameters on TOC and color removal were investigated

• A new operational table was produced on the basis of stoichiometric ratios

• Time-dependant dosing of H2O2 helps to obtain higher efficiency levels

Abstract

The operation conditions for the implementation of the Fenton process are of

utmost importance because there are problems related to the proportion of H2O2

dosage to COD and Fe2+ dosage to the specified H2O2 amount. The relevant lite-

rature shows that COD/H2O2 ratios range between 0.0084 and 113.9. Similarly,

the COD/Fe2+ ratio varies between 0.079 and 292.6, while the H2O2/Fe2+ ratio

varies between 0.09 and 287. Moreover, the ratio of the maximum value to the

minimum value used in the operations on the basis of COD is 13560 for COD/

/H2O2 (0.0084-113.9), 2210 for COD/Fe2+ (0.079-174.7), and finally 3190 for

H2O2/Fe2+ (0,09-287). The aim of this study was to re-evaluate these values that

significantly differ from each other with specific emphasis on textile wastewater

and considering stoichiometric ratios. Results showed that values ranging

between 0.43 and 4.0 for COD/H2O2 and those ranging between 0.75 and 3.0 for

H2O2/Fe2+ are more suitable. The results showed that in a Fenton process con-

ducted by dosing H2O2 at different times, the TOC reduction efficiencies inc-

reased from 80.8 up to 88.9%. Similarly, the color reduction efficiency also rose

from 96.5 to 98.7%.

Keywords: Fenton’s oxidation; operational parameters; oxidation; stoi-chiometric ratio, TOC removal.

Wastewaters are treated physically, chemically,

and biologically. The physical methods use separ-

ation processes rather than a treatment processes. At

the end of these processes, contaminants are gener-

ated in a concentrated manner [1]. At the end of the

biological processes, on the other hand, one of the

contaminants that one emerges is sludge [2]. Even

Correspondence: K. Yetilmezsoy, Department of Environmental

Engineering, Faculty of Civil Engineering, Yildiz Technical Uni-

versity, 34220, Davutpasa, Esenler, Istanbul, Turkey. E-mail: [email protected] * These authors do not have an active academic position at the

Faculty of Civil Engineering, Yildiz Technical University. Paper received: 7 September, 2015 Paper revised: 3 December, 2015 Paper accepted: 8 December, 2015

though the contaminants generated by chemical

methods, such as clarification-type processes, are dif-

ferent, they are still present inside the chemical sludge

[3], whereas the contaminants can be fully avoided by

using advanced treatment processes. The oxidation

processes in particular aid in the complete elimination

of these contaminants. The Fenton process has a

special characteristic among the other oxidation pro-

cesses. The difference of the Fenton process is that it

generates ●OH, which helps introduce the synergistic

effect of peroxide, an oxidant, the iron catalysis [4].

This process is especially efficient for COD parame-

ters that are difficult to reduce, when the BOD/COD

ratio is low. When a Fenton process is applied to a

COD parameter, which cannot be reduced by means

F. ILHAN et al.: EVALUATION OF OPERATIONAL PARAMETERS… Chem. Ind. Chem. Eng. Q. 23 (1) 1119 (2017)

12

of other processes, high efficiencies can be obtained.

The reason for this is that the oxidation capacity of

hydroxyl radicals is higher than the other oxidants [4].

Based on this, one can conclude that the Fenton pro-

cess can be used on a variety of wastewaters such as

those produced by activities related to textiles [5],

leachates [6], paints [7], olive mills [8], pharmaceut-

icals [9], tanneries [10] and surfactants [11]. In addi-

tion, whenever the desired level of efficiency cannot

be obtained by biological refinement, the Fenton pro-

cess can also be used as a final treatment method to

shift COD values to the desired levels [12].

In spite of these advantageous characteristics,

the Fenton process brings about a number of import-

ant economic problems [13]. The major reason behind

this is that normally the peroxide dose and the

amount of iron to be used cannot be finely adjusted.

The addition of peroxide and iron in measurements

that are less than needed causes low efficiency

levels. On the other hand, excessive use of peroxide

creates problems, not only because it requires reduct-

ion (reduction of excessive peroxide), but also it

results in additional costs. Similarly, the addition of

excessive iron also brings about extra costs and also

results in extremely significant operational problems

such as the sludge caused by too much iron. For

these reasons, the selection of operational conditions

for a Fenton process is of utmost importance in terms

of a correct and efficient operation [4].

The aim of this study was to avoid eventual red-

undant and excessive H2O2 and Fe2+ additions, and

hence to prevent the formation of redundant sludge

and excessive peroxide, based on the stoichiometric

ratios that should be ensured by selecting the appro-

priate operational conditions. In addition, a physico-

chemical study was carried out on a textile waste-

water sample, by using less chemicals on the basis of

stoichiometric ratios, and it was aimed to demonstrate

that high reduction efficiencies can actually be

obtained by using low amounts of chemicals.

MATERIALS AND METHODS

Selection of operational parameters

In this study, the effects of the operational con-

ditions such as initial pH value, and COD/H2O2 and

H2O2/Fe2+ ratios on the Fenton process were inves-

tigated. Since the priority of this study was to define

COD/H2O2 and H2O2/Fe2+, these parameters were

investigated in more detail. After defining the optimum

values, an optimization study was carried out on the

basis of pH values. Finally, the effect of adding the

initial peroxide by dosing on the results was analysed.

Also, a comparative analysis between the first 9

studies and the study-set between 29 and 38 will pro-

vide further insight regarding the effect of dosing the

peroxide to add to the results.

Wastewater characteristics

Textile wastewater was selected as the study

case. The textile wastewater samples were taken

from the Akinal textile factory (Cevizlibag, Istanbul,

Turkey) and kept at 4 C. The wastewater parameters

were regularly measured during this study. The main

characteristics of the textile wastewater samples used

in the experiments are shown in Table 1.

Table 1. Characterization of textile wastewater from Akinal tex-

tile factory; SD – standard deviation

Parameter Unit Mean + SD

Chemical oxygen demand (COD) mg/L 1625±40

Five-day biochemical oxygen demand

(BOD5)

mg/L 570±15

BOD5/COD - 0.35

TOC/COD - 0.32

pH - 4.3±0.1

Total organic carbon (TOC) mg/L 524±10

Color Pt-Co 545±10

Conductivity µS/cm 2370±50

Fenton’s oxidation

A stock solution of 10 g/L of Fe2+ was prepared

by dissolving FeSO47H2O (Merck Chemical Corp.) in

distilled water. In addition to iron sulfate reagent, 30%

H2O2 solution (Merck Chemical Corp.) with a density

of 1.11 kg/L was used in the oxidation process. In

each oxidation test, 500 mL of textile wastewater

sample was collected from the textile industry efflu-

ent. In the first step of Fenton’s oxidation process, the

pH of the textile wastewater was adjusted to the

desired value by the addition of 1 M H2SO4 and 1 M

NaOH. During the whole oxidation process, the pH of

samples were also set at the desired value by adding

these reagents (1 M H2SO4 and 1 M NaOH) gradually

in addition to the pre-adjustment of the pH. The

FeSO47H2O and H2O2 solutions were then added to

the effluent sample and conducted for 5 min of rapid

mixing at 120 rpm using a Jar Test Equipment (VELP

Scientifica, FC6S). The effluent sample was then

gently stirred at 10 rpm for 25 min. After the floccul-

ation process, the sample transferred to a graduated

settling column for 30 min of settling. 100 mL of

supernatant sample was then collected for the further

analyses (TOC and color) after the settling process.

In order to prevent interferences in analytical mea-

surements, the pH of collected supernatant sample

F. ILHAN et al.: EVALUATION OF OPERATIONAL PARAMETERS… Chem. Ind. Chem. Eng. Q. .23 (1) 1119 (2017)

13

was increased to about 6.0 by adding 6 M NaOH gra-

dually to precipitate Fe2+ in the form of Fe(OH)3. Fin-

ally, the MnO2 reagent was then added to remove the

residual H2O2 from the collected supernatant [26].

Analytical procedure

The pH of wastewater samples was measured

by a pH meter (WTW series pH 720) and a pH probe

(WTW, pH-Electrode Sentix 41). The color of waste-

water samples was measured with a Hach Lange

spectrofotometer (model: DR 5000) and determined

as platinum-cobalt (Pt-Co) color unit according to

method 120. Electrical conductivity was measured by

using a multimeter instrument (Hach Lange HQ 40D).

The total organic carbon (TOC) was measured by

using a Hach Lange IL 550 TOC/TN analyzer. All

other experimental analyses were performed by the

procedures described in the Standard Methods of

APHA [34]. These parameters were determined by

the procedures described in method numbers of 5220

C (closed reflux, titrimetric method for COD) and 5210

B (5-day BOD test). The deionized water used in the

experiments was supplied from a purification system

(Meck Millipore Direct-Q 3, 5, 8 Ultrapure Water Sys-

tems). The analyses were carried out at least three

times for each sample to assess method precision.

Stability of the oxidation process and components of

wastewater samples were monitored in the Environ-

mental Engineering Laboratory at Yildiz Technical

University in Istanbul, Turkey.


Literature review on the Fenton’s oxidation process

When the data reported in the relevant literature

are investigated, it is seen that the doses selected

both for peroxide and Fe2+ parameters differ greatly

from each other, even when they are proportional to

the COD values. A variety of the most recent studies,

selected among those available in the literature, are

shown in Table 2.

The cells in normal characters in Table 2 are

those directly reported by the mentioned studies. The

cells given in bold characters, on the other hand,

were calculated based on the values reported in

those studies, because each study reported a dif-

ferent ratio (H2O2/Fe2+, COD/H2O2, COD/Fe2+, Fe2+/

/H2O2, H2O2/COD and Fe2+/H2O2). The obtained res-

ults are quite interesting for a number of reasons. As

seen in Table 2, the H2O2 values determined on the

basis of COD are given in terms of concentrations.

Similarly, Fe2+, which are determined based on the

added H2O2 amount in terms of a mole ratio, are also

reported in terms of concentrations. However,

although it is more practical to report these values in

terms of concentrations, this brings about an impor-

Table 2. Comparison of different process typologies on Fenton’s oxidation

COD / mg L–1 pH Time

min

H2O2 dosage

mg/L

Fe2+ dosage

mg/L t / C COD/H2O2 COD/Fe2+ H2O2/Fe2+ Reference

5320 2-9 60 340-15300 280-5600 - 0.35-15.65 0.95-19 0.2-10 [14]

- 2.5-7 5-240 60-450 0-15 - - - 4-45 [15]

3242 2-6 30-120 16950-42375 200-1000 - 0.076-0.19 3.24-16.21 33.9-169.5 [16]

1140 2-8 0-60 1000-6500 150-1000 - 0.175-1.14 1.14-7.6 3.2-21.67 [17]

11987 2-4 90 13185- 52742.8 217-8686 - 0.23-0.91 1.38-55.2 10-100 [18]

314-404 2-5 0-100 - - - 0.126-0.315 0.609-3.125 1.99-9.90 [19]

564 2.5-7.0 0-300 102-510 22.4-100.8 30-60 1.11-5.53 5.6-25.2 0.61-2.43 [5]

176 ± 13.2 3-5 0-180 187-2240 560-2240 - 0.079-0.94 0.079-0.314 1-3 [20]

725 4 0-30 0-100 0-50 - 7.25-72.5 14.5-145 5-10 [21]

25624 0.75-3.75 35-255 225-900 500-4500 - 28.47-113.9 5.69-51.25 0.09-1.35 [22]

1200 5 - 50-300 20-160 - 4-24 7.5-60 0.5-5 [23]

2150-2770 2-5.5 - 19300-57750 222-2196 25-70 0.043- 0.127 1.12-11.08 96-287 [24]

7500-8400 5.4-9.1 0-150 5780-18,020 300-3000 - 0.44-1.37 2.65–26.5 1.93-34 [25]

1750 2-7 60 200-1200 100-1000 - 1.46-8.75 1.75-17.5 0.5-12 [26]

2533 1.5-3.5 30 44000-266000 552-2210 20-40 0.0095-0.058 1.15-4.59 19.91-201.1 [27]

69600-174000 3-5 0-240 1700-5100 204-3050 - 13.65-102 22.82-57 1.67-8.33 [28]

1670 3-7 0-180 15000-200000 13.8-1380 5-20 0.0084-0.111 1.21-121 0-5 [29]

4528 1-7 30-240 100-800 100-800 - 5.66-45.28 5.66-45.28 0.4-5 [30]

11620-85300 3 - 19800-55200 1250-5000 - 0.23-3.40 6.4-25.6 8-22 [31]

575-2271 2.25-2.47 40 39.0-252.9 13.0-84.3 - 2.27-58 6.82-174.69 3.0 [32]


14

tant problem. When these values are proportioned to

COD or as H2O2/Fe2+, they result in values that greatly

differ from each other.

As can be seen from Table 2, the COD/H2O2

ratio ranged between 0.0084 and 113.9. Similarly,

COD/Fe2+ ratios varied between 0.079 and 174.7, and

H2O2/Fe2+ ratios varied between 0.09 and 287. These

intervals are quite high. Since adding less H2O2 than

necessary brings about problems in terms of effi-

ciency, and excessive H2O2 causes problems not only

in efficiency but also in the reduction of excessive

peroxide, these large intervals pose a great challenge

in this oxidation process. For this reason, the ratio of

COD to the amount of H2O2 used should be deter-

mined stoichiometrically [33]. Reactions occur

smoothly by using 10% more than the stoichiometric

ratios. More than this amount should not be used as

this will cause excessive ●OH in water. The stoichio-

metric COD/H2O2 ratios to be used in the Fenton pro-

cesses can be found in the literature [8]. Based on the

stoichiometric ratio [8], it can be stated that for 2.125

g of H2O2 (or 0.0625 mol H2O2), 1 g of COD (or 1 g O2

= 0.03125 mol O2) is needed with a ratio of about

0.471. Considering a 10% margin of safety, it was

determined that the COD/H2O2 ratio should be equal

to 0.428 or higher. This is due to the fact that it is imp-

ortant for economic reasons to evaluate this ratio in

proportions such as 1:1, even 1:0.5 and 1:0.25, rather

than 1:2.125, considering the synergistic effect.

Therefore, it is quite plausible to use values ranging

between 0.428 and 4.0 for the COD/H2O2 ratio. The

COD/H2O2 ratios lower than 0.428 indicate use of

excessive peroxide, which might result in significant

problems from an engineering perspective.

Another important parameter is the H2O2/Fe2+

ratios. When the effect of H2O2 ions is determined (in

those studies where different doses of H2O2 are

added), a single dose is specified and used for the

Fe2+. However, while the specified Fe2+ amounts are

excessive for low doses of H2O2, it is generally insuf-

ficient for high doses of peroxide additions. Although

the higher doses normally induce better results, this is

not completely due to the oxidation effect of ●OH. The

added Fe2+ transform a part of the H2O2 into ●OH,

while the oxidation effect of the remaining H2O2 plays

an important role in the increase of efficiency. More

Fe2+ additions might help transform all H2O2 into ●OH

and therefore increase the oxidation capacity; how-

ever, it is not economically feasible to use the oxid-

ation capacity of peroxide rather than forming more ●OH. For this reason, in order to obtain more correct

results, the ratio of H2O2 to Fe2+ should be kept cons-

tant, rather than fixing the H2O2 amount in Fe2+ opti-

mization or the Fe2+ amount in H2O2 optimization. To

this end, for the first peroxide optimization, a H2O2/

/Fe2+ ratio, which is well-balanced in terms of its stoi-

chiometric value and reported to be successful by the

relevant literature, should be specified and iron

dosing should be carried out based on this ratio.

Then, the most suitable H2O2/Fe2+ ratios should be

defined by keeping the specified optimum H2O2 dose

constant and modifying the Fe2+ amount. When the

literature regarding the H2O2/Fe2+ ratios are inves-

tigated, it is seen that the values reported change

range within a very wide interval. In certain studies

[22] the H2O2/Fe2+ ratio is equal to 0.09, while in some

others [24] this value may rise to 287. The ratio of

these two values is as high as 3190, which clearly

shows that a suitable value for the H2O2/Fe2+ ratio

should be defined by means of optimization studies.

By taking the added Fe2+ ions and the amount of ●OH

formed by H2O2 into account, the stoichiometric ratios

to be used should be defined. The stoichiometric

equation and ratio to be used for H2O2/Fe2+ is given

below:

2 3 •2 2Fe H O Fe OH OH (1)

When this equation is stoichiometrically inves-

tigated, it is seen that for 56 g of Fe2+, 34 g of H2O2 is

needed with a ratio of about 0.607. By also con-

sidering a 10% margin of safety for the addition of

Fe2+, the new ratio can be defined as 0.552. An addi-

ton of Fe2+ greater than the amount resulting in this

ratio will bring about extra costs and also undesired

chemical sludge at the end of the reaction. Therefore,

lower ratios should not be used. It should be kept in

mind that the proportional lower limit for Fe2+ concen-

tration corresponding to the upper limit of the ratio to

use is 0.55. In this study, as can be seen also in

Table 3, the performance of the reaction was eval-

uated by using the lower amounts of Fe2+, such as the

ratios equal to 1 and 2.

Another proportional expression that is repre-

sentative for a Fenton process is COD/Fe2+. As a mat-

ter of fact, these two parameters are not directly inter-

related. This value is specified based on the H2O2

concentration. When the stoichiometric ratios are con-

sidered, the proportional values suggested to obtain a

synergistic effect and less chemical use are as fol-

lows:

Stoichiometric ratio: COD/H2O2/Fe2+: 1/2.125/3.5

Economical ratio: COD/H2O2/Fe2+: 1/2.125–0.25/3.5–0.35

The calculated ratios were obtained by con-

sidering 10% more than the maximum stoichiometric

ratio. The minimum values, on the other hand, were


15

Table 3. Experimental runs and operational parameters

Runs COD/H2O2 H2O2/Fe2+ Initial pH Dosing

1 0.428 0.55 3.00 -

2 0.5 0.55 3.00 -

3 0.6 0.55 3.00 -

4 0.8 0.55 3.00 -

5 1.0 0.55 3.00 -

6 1.2 0.55 3.00 -

7 1.4 0.55 3.00 -

8 1.6 0.55 3.00 -

9 1.8 0.55 3.00 -

10 2.0 0.55 3.00 -

11 0.5 0.55 3.00 -

12 0.5 0.75 3.00 -

13 0.5 1.0 3.00 -

14 0.5 1.2 3.00 -

15 0.5 1.4 3.00 -

16 0.5 1.6 3.00 -

17 0.5 1.7 3.00 -

18 0.5 1.8 3.00 -

19 0.5 1.9 3.00 -

20 0.5 2.0 3.00 -

21 0.5 1.0 2.00 -

22 0.5 1.0 2.25 -

23 0.5 1.0 2.50 -

24 0.5 1.0 2.75 -

25 0.5 1.0 3.25 -

26 0.5 1.0 3.50 -

27 0.5 1.0 3.75 -

28 0.5 1.0 4.00 -

29 0.5 1.0 3.00 1

30 0.5 1.0 3.00 2

31 0.5 1.0 3.00 3

32 0.5 1.0 3.00 4

33 0.5 1.0 3.00 5

34 0.5 1.0 3.00 10

35 0.5 1.0 3.00 15

36 0.5 1.0 3.00 30

37 0.5 1.0 3.00 45

38 0.5 1.0 3.00 60

39 0.5 1.0 3.00 60

40 0.5 1.0 3.00 60

41 0.5 1.0 3.00 60

42 0.5 1.0 3.00 60

43 0.5 1.0 3.00 60

defined as one tenth of the maximum value, specified

considering the synergistic effect. Optimization work

to be carried out between these values is thought to

help define more appropriate operational condition,

addressing higher efficiencies and lower chemical

consumptions. The operational conditions specified in

this study were optimized at each stage before pas-

sing to the next one, with the aim to increase the effi-

ciency by using optimum values in the subsequent

study-sets. The obtained results were discussed

under different sub-titles in this section.

Effect of COD/H2O2 ratio

The most important parameter in a Fenton pro-

cess is the hydrogen peroxide ratio. As the H2O2

amount increases, the oxidation capacity and effi-

ciency also increase. However, the dose to apply

should be suitable to the amount of organic contam-

inants in the water. Otherwise, less H2O2 than needed

will reduce the efficiency level, while excessive H2O2

amounts will increase operational costs. Therefore,

the optimization of H2O2 amount is of utmost impor-

tance in the Fenton oxidation process. The color and

TOC reduction efficiencies obtained for different

doses specified on the basis of stoichiometric ratios

are given in Figure 1. The studies were carried out

with a pH value equal to 3.0, as an average of the

interval 2-4. The H2O2/Fe2+ ratio was taken as 0.55,

which is the most suitable value from a stoichiometric

perspective.

When Figure 1 is closely examined, it is seen

that increasing COD/H2O2 ratios (or decreasing per-

oxide amounts) adversely affect efficiency, since

lower peroxide amounts decrease the formation rate

of the ●OH radicals, and consequently lower the oxid-

ation capacity of the process. Therefore, the highest

efficiency level was obtained when COD/H2O2 ratio

was equal to 0.5. When the COD/H2O2 ratio was

taken as 0.5, the resulting TOC and color reduction

values were 80.2 and 96.1%, respectively. Based on

these results, 0.5 was accepted to be the optimum

value for the COD/H2O2 ratio and used for the opti-

mization of H2O2/Fe2+ ratio in the next step.

Effect of H2O2/Fe2+ ratio

In redox reactions, the Fe2+ ratio that helps form ●OH is also as important as the H2O2 in the Fenton

oxidation process. Similarly to the hydrogen peroxide

dosing, the use of higher than necessary Fe2+ doses

will result in additional costs. Using an insufficient

amount of Fe2+, on the other hand, will result in ineffi-

ciently completed redox reactions, which are needed

for the formation of ●OH. Therefore, the Fe2+ doses to

add must be optimized for better results. To this end,

a series of studies were carried out using H2O2/Fe2+

doses, determined in compliance with stoichiometric

ratios. The obtained results are given in Figure 2. In

these studies, the pH was 3.0, which is the average

value for the 2-4 interval. The COD/H2O2 ratio, on the


16

other hand, was 0.5, which was found to be the opti-

mum value.

When the obtained results are investigated, it is

clearly seen that increasing H2O2/Fe2+ ratios (or

decreasing Fe2+ amounts) slightly decrease the color

and TOC reduction levels. Increasing the H2O2/Fe2+

ratios indicates insufficient Fe2+ additions. Conse-

quently, a lower efficiency level can be observed due

to the inhibition of the formation of ●OH. The optimum

value for the H2O2/Fe2+ ratio was found to be 1.0. The

color and TOC reduction values determined by using

these values were 80.8 and 96.2%, respectively.

Effect of pH on TOC and color removal

Like in all chemical reactions, for a Fenton oxid-

ation process to be able to be carried out efficiently,

there is a suitable pH interval. For a Fenton reaction,

it is known that this interval corresponds to the acidic

range (generally between 2.0 and 4.0) [33,35]. An

optimization study on pH level was carried out to

further elaborate on this value between 2.0 and 4.0.

The results obtained from this study are shown in

Figure 3. This optimization was carried out by taking

H2O2/Fe2+ and COD/H2O2 ratios as 1.0 and 0.5,

respectively, which were found to be the optimum

values in previous studies.

When the obtained results are investigated, it is

seen that the highest efficiency levels are obtained

with a pH value equal to 3.0, because maintaining the

pH value constant at 3.0 (also at the previous opti-

mization studies) increased the reliability of the res-

ults obtained from these optimization studies. Under

these circumstances, the color and TOC reduction

efficiencies were found to be 80.8 and 96.2%, res-

pectively. The results indicated that both TOC and

color removal levels were lower, compared to results

obtained when the pH value was kept equal to 3.0.

This could be due to the decrease in the synergistic

effect of H2O2 and Fe2+ [26].

Effect of H2O2 dosing on TOC and color removal

●OH plays an active role in the reaction mech-

anism of the Fenton oxidation process as oxidant. For

the formation of these radicals, the addition of H2O2 –

Figure 1. Effects of COD/H2O2 ratio on color and TOC removal efficiencies (H2O2/Fe2+ = 0.55, pH 3.0, reaction time = 30 min).

Figure 2. Effects of H2O2/Fe2+ ratio on color and TOC removal efficiencies (COD/H2O2 =0.5, pH 3.0, reaction time = 30 min).


17

in one go, or in different time periods during the react-

ion – is of great importance. For this aim, a series of

studies have been conducted regarding different dos-

ing types in photo-Fenton and solar photo-Fenton

procedures [36-37]. In classical Fenton studies, on

the other hand, the addition of single-step H2O2 is

generally preferred [38-40]. At this stage of the study,

the effects of different H2O2 dosing on color and TOC

reduction were attempted to be investigated. To this

end, the single-step method was not preferred and

the dosing of H2O2 addition was carried out in multiple

steps (2 times (per 15 min), 3 times (per 10 min), 4

times (per 7.5 min), 5 times (per 6 min), 6 times (per 5

min), 10 times (per 3 min), 15 times (per 2 min), 30

times (per 1 min), and 45 times (per 40 s)] in 30 s and

in continuous mode.

The obtained results clearly showed that time-

-dependent dosing of H2O2 might help to obtain higher

efficiency levels. It was seen that by carrying out

dosing at an extended period of time, TOC reduction

efficiency rose from 80.8 up to 88.9% for the same

amount of dosing. Similarly, the color reduction effi-

ciency also increased from 96.5 to 98.7% by time-

-dependent dosing of H2O2. The higher occurrence

rate of the reactions by dosing is also reflected in the

efficiency levels. An instantaneous dosing will play

the role of the oxidation capacity of peroxide instead

of turning all the peroxide into ●OH at one time.

Therefore, it is apparent that the dosing of hydrogen

peroxide for an extended period of time will consi-

derably improve the reduction efficiency levels.

CONCLUSIONS

This study aimed to provide the synchronization

between studies carried out on the operation con-

ditions intrinsic to the Fenton process that is effect-

ively used for strong organic compounds. When the

stoichiometric ratios are not considered, the exces-

sive use of chemicals results in not only additional

costs, but also in the formation of undesired amounts

of sludge. Similarly, the use of insufficient amounts of

reactant results in low reduction efficiencies. There-

fore, an evaluation of the most recent studies on the

Fenton process in the relevant literature was made,

and the need for such a study was demonstrated.

From a dosing point of view, it was shown that an

apparent increase in efficiency levels can be obtained

in the Fenton processes. At the same time, synchro-

nization in operational conditions can be provided and

the efficiency of different studies carried out in this

field can be compared to each other more easily. An

equivalent reduction in efficiency levels can be

obtained with lower chemical consumptions when the

stoichiometric ratios are taken into consideration.

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Figure 3. Effects of pH on color and TOC removal efficiencies (COD/H2O2 = 0.5, H2O2/Fe2+ = 1.0, reaction time = 30 min).


18

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Manag. 90 (2009) 410-422

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41 (2006) 109-120

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19

FATIH ILHAN

KAAN YETILMEZSOY

HARUN AKIF KABUK

KUBRA ULUCAN

TAMER COSKUN

BUSRA AKOGLU

Department of Environmental

Engineering, Faculty of Civil

Engineering, Yildiz Technical

University, Davutpasa, Esenler,

Istanbul, Turkey

NAUČNI RAD

PROCENA RADNIH PARAMETARA I NJIHOV ODNOS SA STEHIOMETRIJOM FENTON OKSIDACIJE OTPADNE VODE TEKSTILNE INDUSTRIJE

Radni uslovi Fentonovog procesa su od najveće važnosti jer postoje problemi vezani za

odnos doziranja H2O2 sa HPK i doziranja Fe2+ sa specifičnom količinom H2O2. Relevantna

literatura pokazuje da su vrednosti odnosa HPK/H2O2 u opsegu između 0,0084 i 113,9.

Slično, odnos HPK/Fe2+ je u opsegu između 0,079 i 292,6 dok odnos H2O2/Fe2+ varira iz-

među 0,09 i 287. Pored toga, odnos maksimalne i minimalne vrednosti korišćenih na bazi

HPK ima vrednost 13560 za odnos HPK/H2O2 (u opsegu od 0,0084 do 113,9), 2210 za

odnos HPK/Fe2+ (u opsegu od 0,079 do 174,7) i konačno 3190 za odnos H2O2/Fe2+ (u

opsegu od 0,09 do 287). Cilj ovog rada je da ponovo proceni ove vrednosti koje se zna-

čajno razlikuju jedna od druge sa posebnim naglaskom na otpadne vode tekstilne indus-

trije uz razmatranje stehiometrijskih odnosa. Ovi rezultati pokazuju da su vrednosti u op-

segu od 0,43 do 4,0 za HPK/H2O2 dok su vrednosti u opsegu od 0,75 do 3,0 mnogo

pogodnije za H2O2/Fe2+. Takođe, ovi rezultati pokazuju da doziranje H2O2 u različitim vre-

menima Fentonovog procesa može da poveća efikasnost smanjenja ukupnog organskog

ugljenika od 80,8 do 88,9%. Slično, efikasnost smanjenja obojenosti se povećava sa 96,5

na 98,7%.

Ključne reči: Fentonova oksidacija, radni parametri, oksidacija, stehiometrijski

odnos, uklanjanje ukupnog organskog ugljenika.






21

JAVAD AHMADISHOAR1

S. HAJIR BAHRAMI1

BARAHMAN MOVASSAGH2

SEYED HOSEIN AMIRSHAHI1

MOKHTAR ARAMI1

1Textile Engineering Department,

Amirkabir University of

Technology, Tehran, Iran 2Chemistry Department, K.N. Toosi

University of Technology, Tehran,

Iran

SCIENTIFIC PAPER

UDC 631.82/.85:549.67:553:66

https://doi.org/10.2298/CICEQ150116049A

REMOVAL OF DISPERSE BLUE 56 AND DISPERSE RED 135 DYES FROM AQUEOUS DISPERSIONS BY MODIFIED MONTMORILLONITE NANOCLAY

Article Highlights

• A modified montmorillonite was used for adsorption of two disperse dyes

• Modified nanoclay could effectively adsorb the disperse dyes from their aqueous dis-

persion

• The molecular weight and structure of dyes affected the dye adsorption from their

aqueous dispersion

• The major interactions between the nonionic dyes and organoclay are of hydrophobic

nature

• Modified montmorillonite can be used for dye removal from textile wastewater indus-

tries

Abstract

In this study modified montmorillonite was used as an adsorbent for the removal

of two selected disperse dyes i.e., Disperse Blue 56 (DB) and Disperse Red 135

(DR) from dye dispersions. The adsorption equilibrium data of dyes adsorption

were investigated by using Nernst, Freundlich and Langmuir isotherm models.

The adsorption kinetics was analyzed by using different models including

pseudo-first-order, pseudo-second-order, Elovich and Intraparticle diffusion

model. The Freundlich isotherm was found to be the most appropriate model for

describing the sorption of the dyes on modified nanoclay. The best fit to the

experimental results was obtained by using the pseudo-second-order kinetic

equation, which satisfactorily described the process of dye adsorption. Although

different kinetic models may control the rate of the adsorption process, the

results indicated that the main rate limiting step was the intraparticle diffusion.

The results showed that the proposed modified montmorillonite could be used as

an effective adsorbent for the removal of disperse dyes even from highly

concentrated dispersions.

Keywords: isotherm, kinetic, modified nanoclay, disperse dye, waste water.

The use and application of synthetic dyes in

number of industries such as textile, plastic, paper,

leather and cosmetic has gradually increased in

recent years. The discharge of large amount of col-

ored effluents, produced from these industries, into

water sources leads to water pollution and causes

major environmental problems. The removal of dyes

Correspondence: H. Bahrami, Textile Engineering Department,

Amirkabir University of Technology, Tehran, Iran. E-mail: [email protected] Paper received: 16 January, 2015 Paper revised: 19 November, 2015 Paper accepted: 28 December, 2015

from wastewater is one of the critical issues for the

dye manufactures and consumer industries.

The aromatic rings in the structure of most syn-

thetic dyes make them biologically non-degradable

[1]. A variety of methods could be suggested to eli-

minate dyes from wastewater. Some methods, such

as ozonation and oxidation, decompose [2,3] the dye

molecules to simpler compounds, which are less envi-

ronmentally hazardous. Techniques such as coagul-

ation and adsorption extract the dye molecules from

the solutions, i.e., water.

In general, the most important characteristics of

the adsorption based techniques are their lower cost,

J. AHMADISHOAR et al.: REMOVAL OF DISPERSE BLUE 56… Chem. Ind. Chem. Eng. Q. 23 (1) 2129 (2017)

22

better availability, high efficiency, and the possibility

to treat dyes in higher concentration forms in com-

parison to the other methods [4]. Various materials

have been used as adsorbents for the dye removal,

i.e., bagasse [5] and TiO2 [6]. As described in the lit-

erature, the activated carbon is the most widely used

among all conventional adsorbents. However eco-

nomical issues make it an expensive material for

water treatment applications [7].

Nanoclays are low-cost adsorbents, which have

been widely employed in dye user industries to ads-

orb dyes from wastewater [8]. Sodium montmorillonite

belongs to the smectite group and is composed of 2:1

type layered silicate with two tetrahedral silicates and

one octahedral layer with a considerable cation

exchange capacity in the octahedral layer. The

sodium montmorillonite is the more common form of

the five natural varieties of smectite [9-12]. The main

characteristic of sodium montmorillonite is its high

cation exchange capacity. Due to substitution of cat-

ions in the layers, there is a negative charge in the

octahedral layer which can be compensated by some

cations. Sodium cations in sodium montmorillonite

can be replaced by some chemical components. This

could lead to surface modifications of the clays [12].

These modifications not only change the surface

characteristics of the clays from hydrophobic nature

to hydrophilic one, but also increase the adsorption

capacity of the clays [13].

The literature survey indicated that the use of

such modifications of the clay surface, led to improve-

ment in the adsorption capacity of the different dye

classes [14]. Different chemical components have

been used for surface modification of nanoclays.

Grafting of functional polymers to the surface of the

clays [15], using ionic liquids such as imidozolium,

pyridinium [16,17] and phosphonium derivatives [18],

1,6-diamino hexane [19] and cationic surfactant [20]

are some of the reported modification types on nano-

clays. The quaternary ammonium compounds are

also typical materials used for nanoclay modification.

Modified clays have been used for removal of differ-

ent classes of dyes from wastewater. Zohra et al. [7]

used modified bentonite for removal of direct dyes.

Modified bentonite was used to remove reactive dyes

from waste water [21]. In another report acid dyes

were removed from wastewater using modified bento-

nite [22]. There is another report on removal of cation

dyes from wastewater using modified montmorillonite

and sepiolite [23,24]. To the best of our knowledge,

there have been relatively few works investigating the

removal of nonionic dyestuffs, i.e., disperse dyes,

from wastewater by using nanoclays.

In this study, a modified montmorillonite with an

aromatic quaternary ammonium modification was

used for the removal of aqueous dispersions of two

disperse dyes with different structures and molecular

weights. Some physicochemical aspects of the rem-

oval reaction such as adsorption isotherm, equilibrium

conditions and adsorption kinetics were investigated.


A modified montmorillonite nanoclay (Nanofil

3010, QA-MMN) was purchased from Süd-Chemie

Company. Sodium montmorillonite (Cloisite-Na+,

MMN) with a cation exchange capacity of 92.6

cmol/kg, was obtained from Southern Clay Products,

Inc. Two commercial grade disperse dyes, i.e., Ser-

ilen Blue Rl (Disperse Blue 56 from Yourkshire Chem-

ical) and Sumikaron Red S-GG (Disperse Red 135

from Sumitomo Chemical) were selected as disperse

dyes and denoted as DB and DR, respectively. The

chemical structures of the employed dyes are illus-

trated in Figure 1. The other chemicals, including

sodium hydroxide, hydrochloride acid and ethanol,

were of analytical grade, supplied by Merck and used

without any purification.

NH2

NH2

O

O

OH

OH

Cl

(a)

N

NN

HN

N

O

O

O

O

O

O

O

(b)

Figure 1. Chemical structure of dye: a) Disperse Blue 56,

b) Disperse Red 135.

Adsorption experiment

All the adsorption experiments were carried out

at 24 C with a constant agitation speed of 500 rpm.

Dye dispersions with specified concentrations were

prepared. Then the nanoclays (MMN and QA-MMN)

were added to the separate dispersions and stirred by

a mechanical stirrer. After the adsorption processes


23

the nanoclay was centrifuged for 5 min at 5000 rpm.

The concentration of the residual dye in the bath was

measured using a Carry 100 UV-Vis spectrophoto-

meter. The decrease in the dye concentration in the

dye bath was accepted to be the amount of dye ads-

orbed by the nanoclays. In order to ensure the com-

plete dissolution of the dye dispersion, a 10 volumet-

ric parts of the dispersion was mixed with 90 volu-

metric parts of pure ethanol. The adsorption capacity

(qe) was calculated as follows:

e 0 e( ) /q c c V W (1)

where c0 (g/L) is the initial concentration of the dye, ce

(g/L) is the concentration of the dye at equilibrium, V

(L) is the volume of the dye dispersion, and W (g) is

the mass of the adsorbent. The quantity of the ads-

orbed dye at time t, qt (g/g), was calculated using the

following equation:

0( ) /t tq c c V W (2)

where ct (g/L) is the concentration of the dye at any

time t.

Adsorption Isotherm

The dye adsorption isotherm was investigated

by treating 200 ml dye dispersions having concen-

trations in the range of 0.1-1 g/L at pH 5 with 1 g of

the modified nanoclay. In order to analyze the iso-

therm data, three models, i.e., Nernst, Langmuir and

Freundlich, were employed.

Adsorption kinetic models

Although several adsorption kinetic models were

suggested for such an adsorption process, they

should be confirmed by the mechanism of the dye

adsorption on a sorbent [4]. The adsorption experi-

ments were carried out by treating 100 ml dye dis-

persions with concentration of 1 g/L at pH 5 with 1 g

of the nanoclay for different times. In order to inves-

tigate the dye adsorption kinetics, three adsorption

models were used - intra-particle diffusion, pseudo-

-first-order and pseudo-second-order.


The effect of pH on the adsorption process

The initial pH value of the dye dispersion is an

important parameter, which controls the adsorption

process and affects the surface charges of nanoclay.

The adsorption study for disperse dyes on modified

nanoclay was carried out using 100 ml dye dispersion

with concentration of 0.2 g/L and 0.2 g adsorbent at

24 C and pH in the range of 3–11 for 2 h.

The results showed that the adsorption of the

two disperse dyes on clay depended on the initial pH

of the dye dispersions. It could be seen from Fig. 2

that the amount of dye adsorbed on the nanoclay

increased as the pH of DR and DB decreased from 7

to 3 and from 9 to 3, respectively. At higher values of

pH, a small increase in the adsorbed amount of dyes

was observed. Non-ionic dyes are negatively charged

in solution [25,26] and the modified nanoclay has

positively charged surface [7] especially at lower pH.

The electrostatic interactions between the negatively

charged dye molecules and the positively charged

clay surface resulted in a higher dye removal at lower

pH. At higher pH, the abundance of OH- competing

with the negatively charged dye molecules for the

adsorption sites led to lower adsorption of both dis-

perse dyes on the modified nanoclay. Similar obser-

vations were reported in other studies [19,25,27].

Figure 2. Effect of initial pH on the adsorption of disperse dyes

on modified nanoclays (QA-MMN).

Rate of adsorption

In order to study the adsorption rate and to

ensure complete adsorption of the dyes, the experi-

ments were carried out for more than 4 h. The amount

of the adsorbed dye on two types of clays (g/L) as a

function of the contact time (in minutes) is shown in

Fig. 3. The results showed that the rate of dye ads-

orption on nanoclay in the initial stage of the contact

period was significantly high and the maximum ads-

orption rate was registered in the initial 20 min of the

contact time for both dyes. After the initial stage of the

adsorption process, the rate of dye adsorption gradu-

ally decreased and reached a steady-state value

identified as the equilibrium loading capacity, qe.

Similar trend was reported in other studies [1,27-29].

The sharp slope in the first stage (higher initial

sorption rate) could be attributed to a large number of

vacant adsorption sites on the surface of the clay

available in the initial stage of the adsorption process

[7]. This led to an increase in the dye concentration

gradient between the dispersion and the surface of


24

the adsorbent [7]. As time passed, the surface of the

clay was gradually occupied by dye molecules, which

resulted in less number of sites on the clay surface

accessible to the dye molecules. This led to decrease

in the tendency of the dye towards the clay surface

[1,28]. The reduced adsorption rate showed a pos-

sible presence of a mono-layer of dye molecules on

the nanoclay surface [1,4,29,30], which was evident

for both disperse dyes.

Figure 3. Effect of contact time for adsorption of DB and DR on

MMN and QA-MMN

Figure 3 also showed differences in the dye

adsorption in the first and last stage of the adsorption

process for both disperse dyes. In the first stage of

the adsorption process (the first 20 min), the amount

of DB dye adsorbed on the nanoclay surface was

rather higher than the amount of DR dye. However

after this period, the adsorption behavior of the dyes

changed and the DB dye showed slower adsorption

rate in comparison to DR dye. Consequently at

steady state the amount of the adsorbed DR dye on

the clay surface is higher than the one of DB dye. It

seemed that, the differences could be related to the

molecular structures of the employed dyestuffs. In

fact, the smaller structure and lower molecular weight

of DB led to a higher mobility of the dye during the

first stage of the adsorption process in comparison to

DR and consequently to a higher adsorption rate. In

the next step it seemed that the major contributive

forces between the nonionic dye and organoclay were

the van der Waals forces and hydrophobic interact-

ions [7,14].

The negatively charged surface of the nanoclay

may adsorb the cationic surfactants in two steps.

Firstly via an ion exchange mechanism, a monolayer

of cationic surfactants on the surface of the clay was

formed. The positively charged ends of the cationic

surfactants were exchanged with the interlayer

exchangeable cations of the clay (Na+) and the hyd-

rophobic head of the cationic surfactants was arranged

outward (Fig. 4a). Secondly through hydrophobic-hyd-

rophobic interactions, cationic surfactant alkyl chains

were attached to the outer alkyl chain of the mono-

layer (Fig. 4b) [31]. For the Nanofil 3010 sample, a

diffraction peak at 2θ = 7.65 corresponding to a

basal spacing of 11.56 Å was observed. According to

the calculated basal spacing, it could be ensured that

the monolayer arrangement was formed [32].

In our study, according to the adsorption/par-

tition model, the organic fraction of the surfactant

modified clay, containing a long alkyl chain, behaved

as a partition medium and the partition occured

through interaction of the dyes with the cationic surf-

actant of the modified nanoclay. The cationic surf-

Figure 4. Schematic diagram of the formation of surfactant on the surface of the clay a) monolayer, b) bilayer formation and

c) the interactions of DR molecules with the surfactant modified adsorbent.


25

actants created hydrophobic regions [33]. The hydro-

phobic portion of the adsorbent surface had greater

affinity for dissociated species of dyes in aqueous

solution and the major contributive forces between

the nonionic dye and the organoclay were the Van der

Waals forces and hydrophobic interactions (Figure

4c). The partition of the dyes into hydrophobic regions

plays an important role in the dye uptake.

The dye with the larger molecular structure (DR)

could be adsorbed better than the DB dye on the

nanoclay surface due to the stronger van der Waals

and hydrophobic interactions (interaction between the

phenyl ring of the dyes and the CH2 group of the

modified adsorbent). A good example of a high sorp-

tion of the disperse dyes on the clay surface modified

with quaternary ammonium with aromatic rings was

the sorption of disperse dyes on Cloisite 10A, where

the Van der Waals forces were the main interaction

forces between the aromatic systems [14]. The affinity

of montmorillonite (MMN) toward these dyes from dis-

persion was lower in comparison to the modified

nanoclay (QA-MMN). From the results shown in Fig. 3

it could also be concluded that the times for achieving

the maximum adsorption for DB and DR were 60 and

120 min respectively, which were chosen for the next

experiments in this study.

Adsorption isotherm

The adsorption isotherm indicated the distribut-

ion of dye adsorbed on the nanoclay and dye in bath

in equilibrium state. The isotherm plots for both dyes

are shown in Fig. 5.

Figure 5 showed the experimental and predicted

data for three isotherms. The parameters of these

isotherms for the employed dyes are given in Table 1.

Based on the individual plots and the high regression

correlation coefficient R2, it could be concluded that

the Freundlich isotherm had better fits than the other

isotherms.

Adsorption kinetic models

The determination of the adsorption kinetics can

provide information about the rate of dye adsorption

on the adsorbent surface form dye dispersion as well

as the adsorption mechanism [4,30]. In order to study

the adsorption kinetics in this study, the experimental

data were analyzed using pseudo-first-order, pseudo-

-second-order, Elovich equation and intra-particle dif-

fusion models. The linear regression correlation coef-

ficient, R2, was used as a criterion to select the model,

which gives the best fit to the experimental data.

Figure. 5. Adsorption isotherm of two disperse dyes on

QA-MMN.

Pseudo-first-order model

The pseudo-first-order model is expressed by

Eq. (3):

Table 1. Calculated isotherms parameters

Dye Freundlich Nerst Langmuir

kf / L g-1 n R2 K / mL g–1 R2 qm / mg g–1 kb / L mg–1 R2

DR 135.64 0.872 0.99 126.4 0.979 0.2154 0.942 0.744

DB 128.38 0.9 0.994 127.5 0.991 0.3279 0.592 0.749


26

1 e

d( )

dt

t

qk q q

t (3)

where qt is the amount of the adsorbed dye at time t;

k1 is the rate constant of the pseudo-first-order model

(min-1); t is the time (min) and qe is the adsorption

capacity in equilibrium (mg/g). After definite integ-

ration by applying the conditions: qt = 0 at t = 0 and

qt = qt at t = t, Eq. (3) becomes [34]:

e e 1ln( ) lntq q q k t (4)

The linear relation between the eln( )tq q and t

confirmed the validity of the model while k1 and qe are

the slop and intercept of the proposed line, respect-

ively. The results listed in Table 2, showed that value

of the the correlation coefficient for DR is higher than

the one for DB and the values of qe obtained by Eq.

(4) for both dyes are not in agreement with the real

values (0.08056 and 0.0977 g/g for DB and DR, res-

pectively). The results showed that the experimental

data are not in agreement with the pseudo-first-order

kinetic model and this model cannot fully describe the

adsorption kinetics. In Fig. 6a and b are shown the

pseudo-first-order plots for DB and DR, respectively.

Pseudo-second-order model

The experimental data were also analyzed by

using pseudo-second-order model. The obtained

results showed that the pseudo-second-order model

was suitable for low initial concentration of the dye

solution [34]. Dogan et al. [1, 29] suggested that the

rate of this model depends on the amount of the

adsorbate on the surface of the adsorbent and the

capacity of the adsorbent at equilibrium. The model is

represented by Eq. (5):

22 e

d( )

dt

t

qk q q

t (5)

where k2 is the pseudo-second-order rate constant

(g/(mol min)). By integration at boundary conditions:

qt = 0 at t = 0 and qt = qt at t = t, Eq. (5) becomes [34]:

2

e2 e

1

t

t t

q qk q (6)

Again, the model was confirmed by linear rel-

ation between t/qt and t. As seen from Table 2, the

correlation coefficient values for DR and DB are

0.9995 and 0.9992, respectively and qe for both dyes

are in agreement with the real values (80.56 and

97.72 mg/g for DB and DR, respectively). The plots of

(a) (b)

(c) (d)

Figure 6. Plot of ln(qe - qt) vs. time (t),a: DB, b: DR and Plot of t/qt vs. t, c:DB, d: DR.


27

the pseudo-second-order model for DB and DR are

presented in Fig. 6c and d, respectively.

Clearly, the pseudo-second-order kinetic equa-

tion could satisfactory describe the dye adsorption

and the experimental results were best fitted by using

this model.

Elovich equation

The validity of this equation suggests the pre-

sence of reactions involving chemical adsorption of

adsorbate on the adsorbent [4,35]. Equation (8) rep-

resents the mathematical form of the Elovich equa-

tion:

d

exp( )d

tt

qq

t (7)

where, is the initial adsorption rate (mol/(g min))

and is the desorption constant (g/mol). To simplify

the Elvoich equation, it was assumed that 1t . By

taking this into account and applying the boundary

conditions: qt = 0 at t = 0, the simple form of the

Elovich equation could be expressed as [1 ]:

( )tq Ln Ln t (8)

The straight line plot of qt vs. lnt confirmed the

validity of the Elovich equation as a suitable model

describing the kinetics of the adsorption process. The

correlation coefficients of the plots of qt vs. lnt for DB

and DR are 0.739 and 0.8476, respectively which

indicated that this model is not valid for this system.

Intraparticle diffusion model

Any adsorption process may consist of the fol-

lowing transportation steps: a) diffusion of the adsor-

bate in the surface of the adsorbent, b) intraparticle or

pore diffusion and c) sorption of adsorbate on the

adsorbent [1,4,36]. The intra particle diffusion model

would be valid, if the plot of qt vs. t0.5 showed a linear

relation and the intraparticle diffusion is the rate-limit-

ing step. According to Weber et al. [37], the equation

of intraparticle diffusion could be expressed by Eq.

(9):

1

2t pq K t C (9)

where qt is the amount of the adsorbed dye (mg/g) at

time t, Kp (mg/(g min1/2)) is the rate constant for int-

raparticle diffusion and C is the intercept. The intrapar-

ticle diffusion plots for both dyes are shown in Fig. 7.

The values of the intercept are proportional to

the thickness of the boundary layer and the larger

intercept corresponds to a greater boundary layer

effect [1,38,39]. As shown in Fig. 7 the adsorption

process occurred in two steps and the plots of qt vs.

t0.5 consist of two linear steps with different slopes for

each of the employed dyes. Similar trends were rep-

orted in other studies [1,4,13,35,40]. They suggested

that the sorption process occurred in two steps, i.e.,

surface sorption and intraparticle diffusion [4]. In the

first step, the dye molecules adsorbed on the external

surface of the clays and the rate of this step was

relatively fast. The second step occurred when the

external surface of the clays was saturated by the

dye. In this case, the dye molecules diffused in the

clays and were adsorbed on their internal surface.

Alkan et al. [1,4] suggested that the first step of the

diffusion process could be attributed to the macro-

-pore diffusion indicating boundary layer effect and

Table 2. Kinetics data calculated for adsorption of DB and DR on QA-MMN

Dye qe, exp

mg/g

Pseudo-first order Pseudo-second order Elovich equation

R2 K1102 / min–1 qe,cal / mg g–1 R2 K2 / g mg–1 min–1 qe,cal / mg g–1 α β R2

DB 80.56 0.5371 0.39 11.056 0.9995 0.00889 76.92 1.711011 2.352 0.739

DR 97.72 0.9956 5.9 81.451 0.9992 0.00135 111.11 0.819 13.74 0.847

Figure 7. Intra-particle diffusion plots for adsorption of DR and DB on QA-MMN.


28

the second step of the diffusion process was due to

the intraparticle or pore diffusion and was attributed to

micro-pore diffusion. The rate constants (Kp1 and Kp2)

of the intraparticle diffusion model for the employed

dyes are shown in Table 3.

The results listed in Table 3 showed that the

value of the first diffusion rate parameter (kp1) for DR

was different from the one for DB [3]. As discussed in

the previous paragraph, the first diffusion rate was

correlated to the surface sorption. The stronger

hydrophobic properties of the DR dye in comparison

to the one of DB and its affinity to the hydrophobic

nanoclay were the reasons for the higher adsorption

rate of the DR dye.

The slope of the second linear stage determined

the rate parameter corresponding to the intraparticle

diffusion. The value of this parameter for DR is higher

than the one for DB. The linear regression and the

comparison of the regression coefficients R12 and R2

2

for both dyes showed that the regression for both

dyes is linear but the plots do not pass through the

origin. This result is more obvious for DB. It sug-

gested that the adsorption involved intraparticle dif-

fusion and the diffusion could be accepted as a rate-

-limiting step. However, the diffusion is not the only

rate-controlling step [1,7,25,26,39] and other kinetic

models may also control the rate of adsorption.

CONCLUSIONS

In this study, the adsorption behavior of two

disperse dyes on modified montmorillonite was inves-

tigated. Modified montmorillonite could effectively

adsorb the disperse dyes from their aqueous dis-

persion and could be used for dye removal from tex-

tile wastewater industries. Changing the pH of the dye

dispersion could affect the dye removal from dye dis-

persion. An increase of the pH to values of above 7

led to increase in the adsorption of both disperse dye

on modified nanoclays. The adsorption rate study

showed that the adsorption of dyes on modified clays

involved two stages and the first stage of the ads-

orption process was faster than the second stage for

both dyes. The selected dyes had different molecular

weight and structure. This affected the dye adsorption

process on the absorbent surface. The isotherm

model, which showed the best fit to the experimental

data for both dyes was the Freundlich linear model

(R2 > 0.99). The kinetics studies showed that the

adsorption of the two disperse dyes could be well

defined by pseudo-second-order kinetic equation with

a high correlation coefficient (R2 > 0.99). The results

showed that the adsorption process involved intra-

particle diffusion and the diffusion could be a rate-

limiting step. However, the diffusion was not the only

rate-controlling step and other kinetic models could

also control the rate of adsorption.

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4054-4062

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C1 / mg g–1 R22 Kp2 / mg g–1 min–½ R1

2 Kp1 / mg g–1 min–½ Dye 59.7 0.0165 0 0.877 1.62 DB

28.55 0.711 0.55 0.989 9.49 DR


29

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ation 252 (2010) 88-96.

JAVAD AHMADISHOAR1

S. HAJIR BAHRAMI1

BARAHMAN MOVASSAGH2

SEYED HOSEIN AMIRSHAHI1

MOKHTAR ARAMI1

1Textile Engineering Department,

Amirkabir University of Technology,

Tehran, Iran 2Chemistry Department, K.N. Toosi

University of Technology, Tehran, Iran

NAUČNI RAD

UKLANJANJE DISPERZNIH BOJA PLAVA 56 I CRVENA 135 IZ VODENIH DISPERZIJA POMOĆU MODIFIKOVANE MONTMORILONITNE NANOGLINE

U ovom radu korišcen je modifikovani montmorilonit kao adsorbent za uklanjanje dve

izabrane disperzne boje disperzno plavo 56 (DB) i disperzno crveno 135 (DR) iz disperzija

boja. Podaci za adsorpcionu ravnotežu su analizirani korišćenjem Nernstovog, Frojndliho-

vog i Lengmirovog izotermskog modela. Kinetika adsorpcije je analizirana korišcenjem

modela pseudo-prvog i pseudo-drugog reda, Elovičevog i modela unutrašnje difuzije.

Frojndlihov izotermski model se pokazao kao nabolji za sorpciju boje na modifikovanoj

nano glini. Jednačina kinetike pseudo-drugog reda na zadovoljavajuci način opisuje brzinu

adsorpciju boje i najbolje fituje dobijene eksperimentalne podatke. Na osnovu dobijenih

rezultata može se zaključiti da se brzina adsorpcija može opisati modelom unutrašnje difu-

zije. Ova difuzija se može smatrati kao stupanj koji limitira brzine, ali ne i kao jedini, s

obzirom na to da i ostali modeli mogu kontrolisati brzinu adsorpcije. Dobijeni rezultati uka-

zuju da predloženi modifikovani montmorilonit može biti efikasan adsorbent za uklanjanje

disperznih boja čak i u slučaju jako koncentrovanim disperzijama.

Ključne reči: izoterma, kinetika, modifikovana nano glina, disperzna boja, otpadna

voda.






31

ALIREZA

EBRAHIMINEZHAD1–3

YAHYA BARZEGAR3,4

YOUNES GHASEMI3

AYDIN BERENJIAN5

1Noncommunicable Diseases

Research Centre, Fasa University

of Medical Sciences, Fasa, Iran 2Department of Medical

Biotechnology, School of Medicine,

Fasa University of Medical

Sciences, Fasa, Iran 3Department of Pharmaceutical

Biotechnology, School of

Pharmacy and Pharmaceutical

Sciences Research Centre, Shiraz

University of Medical Sciences,

Shiraz, Iran 4Department of Biochemistry,

Islamic Azad University, Shiraz

Branch, Shiraz, Iran 5School of Engineering, Faculty of

Science and Engineering, The

University of Waikato, Hamilton,

New Zealand

SCIENTIFIC PAPER

UDC 66.098:582.685.2:546.57:615

https://doi.org/10.2298/CICEQ150824002E

GREEN SYNTHESIS AND CHARACTERIZATION OF SILVER NANOPARTICLES USING Alcea rosea FLOWER EXTRACT AS A NEW GENERATION OF ANTIMICROBIALS

Article Highlights

• Synthesis of silver nanoparticles was developed using Alcea rosea flower extract

• AgNO3 concentration, flower extracts quantity, and reaction temperatures were deter-

mined to be significant factors in the AgNPs biosynthesis

• Prepared AgNPs were spherical in shape with 7.2 nm mean particle size

• Oxygen-bearing functional groups in biochemical compounds from A. rosea were res-

ponsible for reduction of Ag+

• The MIC for AgNPs against E. coli and S. aureus was determined to be 37.5 µg/mL

Abstract

Green synthesis of silver nanoparticles (AgNPs) was developed by treating Ag+

with Alcea rosea flower extract. AgNO3 concentration, flower extract quantity,

and reaction temperature were found to be significant factors in the bioreduction

reaction. Synthesized AgNPs were almost spherical in shape with an average

diameter of 7.2 nm. Fourier transform infrared spectroscopy (FTIR) analysis

revealed that oxygen-bearing functional groups in the A. rosea flower extract are

responsible for reduction of Ag+. The minimum inhibitory concentration (MIC) of

AgNPs against a Gram-positive (Staphylococcus aureus) and Gram-negative

(Escherichia coli) bacteria was determined to be 37.5 µg/ml.

Keywords: Alcea rosea; biochemical reduction; biosynthesis; green syn-thesis; Ag nanoparticles.

AgNPs have found widespread technological

applications due to their unique physicochemical,

optical and catalytic properties [1,2]. There is an inc-

reasing interest for application of AgNPs in house-

holds, medicine and industry [3]. Nowadays, there are

a lot of commercially available products containing

AgNPs, ranging from burn treating materials to anti-

microbial fabrics and paints [3]. The increasing applic-

ation of AgNPs will lead to increased demand for

AgNPs production. To date, various chemical and

Correspondence: A. Berenjian, School of Engineering, Faculty

of Science and Engineering, The University of Waikato, Hamil-

ton, New Zealand. E-mail: [email protected] Paper received: 24 August, 2015 Paper revised: 30 November, 2015 Paper accepted: 20 January, 2016

physicochemical techniques have been used for the

production of AgNPs [4,5]. However, all these

methods suffer from high energy consumption and

the use of toxic chemicals which are potentially dan-

gerous to the environment and human health. There-

fore, there is a need for development of reliable and

green process of AgNPs synthesis.

Green chemistry has emerged as a new concept

for development and implementation of chemical pro-

cesses in order to reduce or eliminate the use of

hazardous substances. Chemically synthesized nano-

particles are not colloidal or physicochemically stable

in aqueous media and therefore capping agents must

be used to increase the particle stability. In biosyn-

thesis reactions, biochemical species attach to the

surface of nanoparticles and act as capping and stab-

ilizing agent in a one-pot reaction [6-12]. Bioactive

A. EBRAHIMINEZHAD et al.: GREEN SYNTHESIS… Chem. Ind. Chem. Eng. Q. 23 (1) 3137 (2017)

32

compounds from microorganisms and plants have a

valuable capability for reduction and capping of

AgNPs without the use of any toxic chemicals and

harsh reaction conditions [6-8,10]. Carbohydrates and

proteins from microbial cells can be effective in red-

uction of Ag+ [12]. In comparison to usage of plant

extracts, biosynthesis of AgNPs using microorg-

anisms needs an elaborated process of culturing and

maintaining microbial cells that, in some cases, could

be pathogenic to humans. The use of plant extract

has advantages such as ease of handling, availability

and a broad viability of metabolites. AgNPs have

been synthesized using leaf extract of various plants

such as black tea, Lippia citriodora (Lemon Verbena),

maple (Acer sp.) and eucalyptus [7,8,13,14]. Other

parts of plants have also been used such as Piper

longum and Crataegus douglasii fruit extract, coffee

powder extract, Nephelium lappaceum, orange peel

extract, oil of Plukenetia volubilis L., Chrysanthemum

morifolium Ramat extract, Medicago sativa and Ster-

culia foetida seed exudate, sorghum bran extract and

Cinnamon zeylanicum bark extract [6,7,10,11,15-21].

Alcea rosea (Althaea rosea) is an important

medicinal herb in many countries. It was used trad-

itionally as expectorant, cooling, diuretic and emme-

nagogue substance. Alcea rosea flowers extract is

used as an anti-inflammatory, febrifuge, demulcent

and astringent agent. Flowers as well as their roots

are used in the treatment of inflammation of the kid-

neys and the uterus. Alcea rosea contains high mole-

cular weight acidic polysaccharides (1.3 to 1.6 million

Dalton) known as mucilages which are abundant in

flowers and leaves. These mucilages are composed

of glucoronic acid, galacturonic acid, rhamnose and

galactose. It also contains proteins, alkaloids and fla-

vonoids [22]. All of these biochemical compounds are

reported to be effective in bioreduction of Ag+ [6-8,10].

According to our best knowledge there is no report on

biosynthesis of AgNPs using Alcea rosea. Therefore,

the current research work aims to investigate: i) the

potential of Alcea rosea flower aqueous extract for

biosynthesis of AgNPs and ii) the antimicrobial effect

of the prepared AgNPs.

EXPERIMENTAL

Materials

Silver nitrate was purchased from Merck. All glass-

ware have been acid washed and then rinsed with

deionised water. All the solutions were prepared using

deionized-Millipore water (resistance >18 M cm).

Preparation of flower extract

Dried flowers of Alcea rosea were initially

washed in deionized water to remove the soil and

dust particles. The aqueous extract was consequently

prepared by mixing 2.5 g of dried flowers with 100 mL

of deionized water in a 250 mL Erlenmeyer flask. The

prepared mixture was boiled for 15 min, then filtered

through Whatman filter paper (Reeve angel®, grade

201) and stored at -20 C.

Preparation of AgNO3 solutions

Solutions of AgNO3 were prepared at 100, 50

and 10 mM concentration. For the initial/stock con-

centration, 1.7 g AgNO3 was dissolved in 100 mL

deionized water to obtain 100 mM solution. Solutions

with lower concentrations (50 and 10 mM) were pre-

pared by two and 10-fold dilutions, respectively.

Desired AgNO3 concentrations in the bioreduction

reactions were achieved by adding 1 mL of corres-

ponding solution to the reaction mixture.

Biosynthesis and characterization of AgNPs

Alcea rosea flowers extract was used as a

source for reducing and capping agent for synthesis

and stabilization of AgNPs in a simplified one-pot

reaction. The impact of various parameters such as

the amount of Alcea rosea flower extract, AgNO3 con-

centration and the reaction temperature was evalu-

ated by conducting several sets of experiments. The

AgNO3 concentration was tested at 1, 5 and 10 mM in

10 mL total reaction volume containing 4 mL (40

vol.%) flower extract at room temperature (27 C).

Impact of various flower extract amounts was also

investigated in the range from 10 to 70% of total

reaction volume at room temperature and 5 mM

AgNO3. The effect of reaction temperature on silver

ions reduction was evaluated at 15, 28, 50 and 75 C

using 5 mM AgNO3 and 4 mL flower extract. All the

reactions were monitored for 24 h.

The optical properties of the produced particles

were analyzed by ultraviolet and visible absorption

spectroscopy (T80+ UV/Vis spectrometer, PG Instru-

ments Ltd.) operated at a resolution of 1 nm within the

range of 300-700 nm. In each analysis, 0.1 mL of the

sample was diluted to 1 mL with deionized water [11].

Further characterizations were done by Transmission

Electron Microscopy (TEM, Philips, CM 10; HT 100

kV), Fourier-transform infrared spectroscopy (FTIR,

Bruker, Vertex 70, FT-IR spectrometer) and X-ray

powder diffraction (XRD, Siemens D5000).

Antimicrobial assay

Escherichia coli PTCC 1399 (ATCC 25922) and

Staphylococcus aureus PTCC 1112 (ATCC 6538)


33

were purchased from Persian Type Culture Collection

(PTCC). Minimum inhibitory concentration (MIC) was

determined using standard microdilution method

(CLSI M07-A8) [23]. In the experiment, AgNPs sus-

pension was prepared in Mueller-Hinton broth (MHB).

To prepare bacterial suspensions for inoculation, bac-

terial cells were cultured in MHB up to turbidity of the

BaSO4 0.5 McFarland standard (OD600 0.11). Then

the 0.5 McFarland suspensions were diluted to 1:20.

Finally, 10 µL of the prepared inoculums’ suspension

was transferred to each well in the 96-well plate con-

taining 90 µL MHB media with AgNPs. For blank

wells, 10 µL of fresh MHB was added to 90 µL MHB

with AgNPs. After 24 h incubation at 37 C, the

OD600 was measured by microplate reader (BioTek,

Power Wave XS2).


Biosynthesis of AgNPs

Reduction of Ag+ to AgNPs resulted in color

change of the reaction solution due to excitation of

surface plasmon resonance (SPR) in the AgNPs.

AgNPs have a typical surface plasmon band absorp-

tion at about 400-450 nm. The UV-Vis spectroscopy,

therefore, can be used as an indirect method to

examine the formation and to some extend charac-

terisation of AgNPs [8,11]. The UV-Vis spectra of the

prepared particles in various concentrations of silver

nitrate are shown in Fig. 1. Increasing the silver nit-

rate concentration from 1 to 5 mM resulted in a major

increase of AgNPs content as indicated by the hyper-

chromic shift in SPR band. No significant increase in

AgNPs concentration was observed when using

higher AgNO3 concentration (>5 mM).

Figure 1. UV-Vis spectra of AgNPs prepared at various

concentrations of silver nitrate: a) 1, b) 5 and c) 10 mM.

SPR bands of the prepared AgNPs using differ-

ent amounts of flower extracts are shown in Fig. 2. By

increasing the flower extract up to 40% of the total

reaction volume, an obvious hyperchromic shift was

observed. However, further increase in the flower ext-

ract content, above 40 vol.% resulted in a significant

hypochromic shift in SPR band. Synthesis of the

metal nanoparticles was conducted in two main steps

namely, a) nucleation and b) growth of nanoparticles.

Organic compounds presence in the reaction mixture

has an inhibitory effect on the particle growth [24-26].

Thus, an optimal value of the Alcea rosea flower ext-

ract is required for reduction of Ag+ to AgNPs.

Figure 2. UV-Vis spectra of AgNPs prepared in various amounts

of Alcea rosea flower extract: a) 10, b) 20, c) 40 and d) 70% of

the reaction volume.

By increasing the volume of flower extract to

more than 40 vol.% of the total reaction volume, a

second absorption peak appeared at about 660 nm.

As shown in TEM micrographs (Fig. 3b), appearance

of this second peak is due to the formation of the

second population of large particles.

Figure 3. TEM micrographs of the prepared AgNPs at various

amounts of flower extract: a) 40 and b) 70 vol.% of the total

reaction volume.

The UV-Vis spectra of the prepared particles at

different temperatures are shown in Fig. 4. Increasing


34

the reaction temperature resulted in significant inc-

rease of the AgNPs concentration, as shown by the

hyperchromic shift in the SPR band. However, the

reaction temperature increase resulted in the appear-

ance of a shoulder in UV-Vis absorption spectra,

indicating the formation of poly-disperse AgNPs [12].

According to the results, room temperature is the

optimal temperature for bioreduction of AgNPs by

Alcea rosea flower extract. Conducting reaction in the

ambient condition could considerably reduce the

energy cost, which is one of the most important

issues in the scale-up process. A flowchart diagram of

the biosynthesis process is illustrated in Fig. 5.

Figure 4: UV-Vis spectra of the prepared AgNPs at various

temperatures: a) 27, b) 40, c) 50 and d) 70 C.

Figure 5. The flowchart diagram of the biosynthesis process.

Characterizations of AgNPs

Particle size distribution was determined by

measuring diameters of one hundred nanoparticles

randomly selected on the TEM images [27]. As shown

in Fig 6, the prepared particles were spherical in

shape with the average diameter of 7.2 nm. The pre-

pared particles were spherical in shape with the aver-

age diameter of 7.2 nm. The crystallinity of the par-

ticles was evaluated by X-ray powder diffraction pat-

terns (XRD, Siemens D5000) using drop coated films

on a glass slide [6,9]. As shown in Fig. 7, four main

characteristic diffraction peaks for silver were obs-

erved at 2θ values 38.2, 44.4, 64.7 and 77.4° due to

reflection from the crystal facets of (111), (200), (220)

and (311), respectively (JCPDS, silver file No. 04-

-0783) [11,13]. Three peaks around 2θ = 32° are indi-

cated by asterisks. Some researchers have attributed

these peaks to the interaction of silver nitrate with

biologic matrixes [28,29].

Figure 6. Particles size distribution of the prepared AgNPs.

Figure 7. XRD pattern of AgNPs indicating four main charact-

eristic peaks for silver, the peaks indicated by asterisk are from

mineral complexes.

FTIR spectra of AgNPs and Alcea rosea flower

extract are depicted in Fig. 8. The bands at 1059 and

1261 cm-1 are from C−O and C−C stretching vib-

rations, respectively. The peak with medium intensity

at 1421 cm-1 could be due to C−H bending vibrations.

Stretching vibrations of aliphatic C−H absorbed IR

radiation at about 2925 cm-1. The absorption peak

from carbonyl groups appeared at 1630 cm-1.

The broad absorption peak of hydrogen bonds

from O−H groups can be seen at 3394 cm-1 which

could overlap with the absorption from N−H bonds

[23,30-32]. Similarity to the flower extract, AgNPs

FTIR spectra indicates that AgNPs are capped with

biochemical compounds from Alcea rosea flower

extract. As it could be observed in the FTIR spectrum,

the main peaks come from oxygen-bearing functional

groups. It is widely believed that oxygen-containing

functionalities are necessary for anchoring of the

metal nanoparticles, and silver ions could easily oxid-

ize these groups [25].


35

Figure 8. FTIR spectra of AgNPs (a) and Alcea rosea flower

extract (b).

Antibacterial assay

Prepared nanoparticles have shown intense

effect on the bacterial growth (Fig. 9). The MIC con-

centrations for S. aureus and E. coli were determined

to be 37.5 µg/ml, which is acceptable compared to the

previously reported concentrations, 10-60 µg/ml [33-

–35]. Silver ions and silver based compounds have

strong antimicrobial effects and have been used for

decades as antimicrobial agents in various fields

[36,37]. AgNPs provide high fraction of exposed

atoms due to their extremely small size and thus

expand the contact surface of silver with microorg-

anisms. It has been confirmed that antimicrobial pro-

perties of AgNPs are due to oxidation of the exposed

silver atoms and release of Ag+ from the surface of

AgNPs [37]. Exposure to the air promoted Ag+ release

and resulted in 2.3-fold increase in the AgNPs anti-

microbial effects [36]. Silver ions are potent oxidants

and can destroy variety of cellular structures. The Ag+

enter into the bacterial cells by penetrating through

the cell wall and consequently turn the DNA molecule

into condensed form which results in the cell death. In

addition, it was also shown that Ag+ binds to funct-

ional groups of proteins, resulting in protein denatur-

ation [35,38]. Metal nanoparticles and particularly

AgNPs are able to destroy the permeability of the

bacterial membranes [34,39]. Exposure of bacterial

membranes to the AgNPs resulted in the leakage of

reducing sugars and proteins and induced the res-

piratory chain dehydrogenases into inactive state [34].

Commonly used antibiotics act very specifically

and target exact physiological points in the microorg-

anisms. This precise strategy provides a chance for

some mutant strains to escape and distribute. To

alleviate this condition and reduce the probability of

new resistant strain appearance, multidrug therapies

have been developed and used; however, this stra-

tegy is ineffective against multidrug resistant strains.

Interestingly, released Ag+ from AgNPs with multi-tar-

geting antimicrobial mechanism of action, significantly

reduced the chance for mutation and development of

a bacterial resistance mechanism. In addition, AgNPs

could increase the potency of the common antibiotics

[33,40-42].

Figure 9. Antimicrobial effect of AgNPs against E. coli

and S. aureus.

CONCLUSION

Alcea rosea flower extract contains bioactive

compounds that are effective in bioreduction of Ag+.

These biochemical compounds contain oxygen bear-

ing functional groups, which act as an anchor for Ag+.

Subsequently, produced AgNPs are capped with hyd-

rophilic biochemical compounds, which make them

colloidally stable. Reaction conditions such as silver

precursor concentration, amount of flower extract and

reaction temperature are the key factors in prepar-

ation of quality-based AgNPs. Therefore, these fac-

tors should be controlled in order to produce uniform

particles with narrow particle size distribution.

Acknowledgments

This investigation was financially supported by

The Fasa University of Medical Sciences, Fasa, Iran,

and The University of Waikato, Hamilton, New Zeal-

and.

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37

ALIREZA EBRAHIMINEZHAD1–3

YAHYA BARZEGAR3,4

YOUNES GHASEMI3

AYDIN BERENJIAN5

1Noncommunicable Diseases

Research Centre, Fasa University of

Medical Sciences, Fasa, Iran 2Department of Medical Biotechnology,

School of Medicine, Fasa University of

Medical Sciences, Fasa, Iran 3Department of Pharmaceutical

Biotechnology, School of Pharmacy

and Pharmaceutical Sciences

Research Centre, Shiraz University of

Medical Sciences, Shiraz, Iran 4Department of Biochemistry, Islamic

Azad University, Shiraz Branch, Shiraz,

Iran 5School of Engineering, Faculty of Sci-

ence and Engineering, The University

of Waikato, Hamilton, New Zealand

NAUČNI RAD

ZELENA SINTEZA I KARAKTERIZACIJA NANOČESTICA SREBRA POMOĆU EKSTRAKTA CVETA Alcea rosea KAO ANTIMIKROBNOG SREDSTVA NOVE GENERACIJE

U radu je razvijena zelena sinteza nanočestica srebra (AgNP) tretiranjem Ag+ ekstraktom

cveta Alcea rosea. Utvrđeno je da su koncentracija AgNO3, količina ekstrakta cveta i

temperatura reakcije značajni faktori ove bioredukcione reakcije. Dobijene AgNP čestice

su sfernog oblika sa prosečnim prečnikom od 7,2 nm. FTIR analiza je pokazala da su za

redukciju Ag+ odgovorne funkcionalne grupe sa kiseonikom prisutne u ekstraktu cveta A.

rosea. Određena je minimalna inhibitorna koncentracija (MIC) AgNP za Gram-pozitivne

(Staphylococcus aureus) i Gram-negativne (Escherichia coli) bakterije i ona iznosi 37,5

µg/ml.

Ključne reči: Alcea rosea, biohemijska redukcija, biosinteza, zelena sinteza, Ag

nanočestice.






39

ALEKSANDRA MIŠAN1

BOJANA ŠARIĆ1

IVAN MILOVANOVIĆ1

PAVLE JOVANOV1

IVANA SEDEJ1

VANJA TADIĆ2

ANAMARIJA MANDIĆ1

MARIJANA SAKAČ1

1Institute of Food Technology in

Novi Sad (FINS), University of Novi

Sad, Novi Sad, Serbia 2Institute for Medicinal Plant

Research “Dr Josif Pančić”,

Belgrade, Serbia

SCIENTIFIC PAPER

UDC 633.12:66.061:616:615

https://doi.org/10.2298/CICEQ150704004M

PHENOLIC PROFILE AND ANTIOXIDANT PROPERTIES OF DRIED BUCKWHEAT LEAF AND FLOWER EXTRACTS

Article Highlights

• Buckwheat leaf and flower extracts were obtained by different extraction procedures

• Rutin and chlorogenic acid were identified as most abundant phenolic compounds

• Some of the extracts were as efficient as BHT in β-carotene bleaching test

• The extracts demonstrated strong ability to inhibit the destruction of erythrocytes

• The extracts prolonged the beginning of the oxidation process in sunflower oil

Abstract

Due to a high content of rutin (2-10%), dried buckwheat leaf and flower (DBLF)

formulations were shown to be efficient in the treatment of vascular diseases. In

order to find a cost effective way for the extraction of antioxidants, the effects of

ethanol/water ratio and temperature on the extraction efficiency of phenolic

compounds and the mechanisms of antioxidant action of the extracts were

tested. Extraction with ethanol/water mixture (80:20, v/v) for 24 h at room tempe-

rature, after the mixture was just brought to boil was demonstrated to be an

efficient and cheap way for obtaining a high yield of rutin (49.94±0.623 mg/g

DBLF). The most abundant phenolic compounds in DBLF extracts were rutin and

chlorogenic acid. Flavonoids, especially rutin, were shown to be the most res-

ponsible for the antioxidant activity in all investigated lipid model systems, acting

as free radical scavengers, electron-donating substances and chelators of iron

ions. In β-carotene bleaching tests, the extracts with the highest activity were as

efficient as BHT (butylated hydroxytoluole). Regarding the results of antihemo-

lytic and Schaal oven tests, the extracts demonstrated remarkable ability to

inhibit the oxidative destruction of erythrocytes and to prolong the beginning of

the oxidation process in sunflower oil.

Keywords: dried buckwheat leaf and flower, rutin, antioxidant activity, lipid oxidation, extraction.

Common buckwheat (Fagopyrum esculentum

Moench) is a highly nutritious pseudocereal known as

a very rich source of antioxidants, especially rutin [1].

The nutritive profile and antioxidant potential of buck-

wheat seeds has been extensively investigated [2-4],

and buckwheat seeds can now be regarded a “func-

tional food” [5]. However, recent studies indicate that

the highest concentration of rutin, up to 10%, is accu-

Correspondence: A. Mišan, Institute of Food Technology in Novi

Sad, University of Novi Sad, Bulevar cara Lazara 1, 21000 Novi

Sad, Serbia. E-mail: [email protected] Paper received: 4 July, 2015 Paper revised: 3 February, 2016 Paper accepted: 12 February, 2016

mulated in leaves and blossoms of the buckwheat

plant [1,6].

Plant phenolic compounds are well known as

highly effective free radical scavengers and antioxi-

dants [7]. The role of an antioxidant in a food product

is related to its ability to inhibit or stop rancidity and/or

deterioration of the nutritional quality [8]. Also, many

natural antioxidants are supposed to have protective

effects against chronic diseases, mainly by scaveng-

ing oxygen radicals, which can deteriorate biological

membranes.

Fagopyri herba is a herbal drug derived from

dried areal tissues of common buckwheat (Fagopy-

rum esculentum Moench) and has been used in the

treatment of vascular diseases [9]. Rutin (quercetin-3-

A. MIŠAN et al.: PHENOLIC PROFILE AND ANTIOXIDANT PROPERTIES… Chem. Ind. Chem. Eng. Q. 23 (1) 3947 (2017)

40

-O-rutinoside), the dominant flavonol glycoside in

DBLF, has been reported to possess antioxidant act-

ivity, to antagonize the increase of capillary fragility

associated with haemorrhagic disease, to reduce high

blood pressure [10], to decrease the permeability of

the blood vessels, to have an anti-oedema effect, and

to reduce the risk of atherosclerosis [11].

Since DBLF can be considered as an ingredient

for designing functional food products, the objective of

this research was to find a cost effective way for the

extraction of phenolic compounds from DBLF in order

to obtain the highest antioxidant activity of the ext-

racts. Water and ethanol/water mixtures were chosen

as nontoxic and environmentally friendly solvents,

which have been shown to be effective in the extract-

ion of quercetin glycosides [12]. Besides the phenol

profiling, the aim of this work was to test the mech-

anisms of antioxidant action of the obtained extracts

in chosen model systems to be able to predict their

possible use.

EXPERIMENTAL

Materials

DBLF (dried areal parts of Fagopyrum escul-

entum Moench collected during the flowering season)

was obtained from the Institute for Medicinal Plants

Research “Dr Josif Pančić” (Belgrade, Serbia) where

a herbarium voucher specimen (No. 31210911) was

deposited. For replicates, three packages of herbal

drug were provided, each containing 500 g of mat-

erial.

Butylated hydroxytoluole (BHT), 1,1-diphenyl-2-

-picrylhydrazyl (DPPH), ethylene diamine tetraacetic

acid disodium salt dihydrate (EDTA), 3-(2-pyridyl)-5-

-6-bis(4-phenyl-sulfonic acid)-1,2,4-triazine (ferro-

zine), ferrous sulfate heptahydrate, linoleic acid

(99%), potassium ferricyanide, sodium carbonate,

Tween 40, trichloracetic acid (TCA), Folin-Ciocalteu's

reagent, standard substances including gallic acid,

protocatechuic acid, caffeic acid, vanillic acid, chloro-

genic acid, syringic acid, ferulic acid, rutin, myricetin,

rosmarinic acid, trans cinnamic acid, naringenin, lut-

eolin, kaempferol, and apigenin were obtained from

Sigma (Sigma-Aldrich GmbH, Sternheim, Germany).

Quercetin was a product of J. T. Baker (Deventer, the

Netherlands), while high-performance liquid chro-

matography (HPLC) grade methanol, formic acid

(HPLC grade) and ethanol 96% were purchased from

Merck (Darmstadt, Germany). Water was purified

using Millipore Elix 10 UV water purification system

(Molsheim, France), and ultrapure water used for

HPLC mobile phase preparation was obtained using

Simplicity UV, Millipore (Molsheim, France).

Preparation of extracts

DBLF (2 g) was mixed either with 50 mL of

water or ethanol/water mixtures (50:50 and 80:20,

V/V). Maceration was performed for 24 h at room

temperature followed by extraction in an ultrasonic

bath (10 min at room temperature). Corresponding

extracts brought to boil before the maceration were

also prepared. The extracts were filtered through the

filter paper (Whatman, Grade 4 Chr, UK) and stored

at -4 C (up to two days) until further use. Preparation

of the extracts for HPLC-MS/MS analyses included

additional evaporation to dryness and redissolving in

an appropriate mobile phase.

Total flavonoid content

Colorimetric aluminum chloride method was

used for determination of total flavonoid content [13],

which is based on the formation of a complex flavo-

noid-aluminium. The probes were prepared by mixing

5 mL of extract, 1 mL of distilled water, and 2.5 mL of

AlCl3 solution (26.6 mg AlCl36H2O and 80 mg

CH3COONa dissolved in 20 mL distilled water). A

blank probe was prepared by replacing AlCl3 solution

with distilled water. The absorbance of probes and

blank probe were measured immediately at 430 nm.

Rutin was used as a standard and results were exp-

ressed as rutin equivalents (RE, g RE per 100 g of

sample). Absorption readings at 415 nm were taken

against a blank sample (reaction mixture without

AlCl3). Rutin (2.5 to 50 μg/mL) was used for the calib-

ration curve construction.

Identification of phenolic compounds by LC-MS/MS

Rapid resolution liquid chromatography with

mass selective detection was performed using an

Agilent Technologies 1200 Series liquid chromato-

graph coupled with Agilent Technologies 6410 triple-

quadropole (QQQ) mass spectrometer. An Eclipse

XDB-C18, 1.8 μm, 4.6 mm50 mm column (Agilent)

was used for separation of 18 phenolic compounds.

The solvent gradient program was created by varying

the proportion of solvent A (0.1 vol.% formic acid in

water) and solvent B (methanol). The following grad-

ient mode was used for phenolic acids identification:

initial 10% B; 0-2 min, 10-43% B; 2-7 min, 43% B;

7-9.5 min, 43-100% B; 9.5-10.5 min, 100% B with flow

rate of 1 mL/min. Identification of flavonoids was per-

formed using the following gradient mode: initial 4%

B; 0-1.5 min, 4-4.5% B; 1.5-4 min, 4.5-10% B; 4-11

min, 10% B; 11-18 min, 10-100% B, with flow rate of

1.2 mL/min. Injection volume was 1 μL. The eluted


41

components were ionized in negative electrospray

ionization (ESI) mode, using nitrogen as nebulizer

(pressure of 50 psi) and drying gas (temperature of

350 C, flow 8 L/min). Extracts used for LC–MS/MS

quantification were dissolved in starting mobile phase

solvent to the concentration of 0.2 mg/mL. All used

standards were dissolved in methanol to prepare

stock solutions of 1 mg/mL and the mix of stock sol-

utions was prepared, with concentration of each com-

pound being 100 μg/mL. Analyses were performed in

MRM (multiple reaction monitoring) mode. Com-

pound-specific, optimized MS/MS parameters are

given in Table 1.

Quantification of phenolic compounds by HPLC-DAD

A liquid chromatograph (Agilent 1200 series),

equipped with a DAD detector and an Eclipse XDB-

-C18, 1.8 μm, 4.6 mm50 mm column (Agilent) was

used for quantification of identified phenolic com-

pounds in DBLF extracts. A single rapid resolution

HPLC method suitable for the determination of 17

phenolic compounds, developed and validated as

previously reported by Mišan et al. [14], was used. In

brief, solvent gradient was performed by varying the

proportion of solvent A (methanol) to solvent B (1

vol.% formic acid in water) as follows: initial 10% A;

0-10 min, 10-25 % A; 10-20 min, 25-60 % A; 20-30

min, 60-70 % A at a flow-rate of 1 mL/min. The total

running time and post-running time were 45 and 10

min, respectively. The column temperature was 30

C. The injected volume of samples and standards

was 5 μL and it was done automatically using an

autosampler. The spectra were acquired in the range

of 210-400 nm and chromatograms plotted at 280,

330 and 350 nm with a bandwidth of 4 nm, and with

reference wavelength/bandwidth of 500/100 nm.

DPPH radical scavenging activity

A modified method of Hatano et al. [15] was

used to determine effect of different extracts on sca-

venging of 1,1-diphenyl-2-picrylhydrazyl (DPPH•) rad-

icals. 0.1 mL of examined extract previously diluted to

obtain at least four different concentrations (0.1 to 5

mg plant material/mL), 1.0 mL of DPPH• (90 μmol/L)

and 2.9 mL of methanol were shaken vigorously and

left to stand in the dark for 60 min. The absorbance

was measured at 515 nm (Cintra 101, GBC scientific,

UV/Vis) against the control (above mentioned mixture

without extract). Results were expressed as the con-

centration (mg plant material/mL) of the extract lead-

ing to 50% reduction of the initial DPPH• concen-

tration (IC50 value). BHT in the concentration range

0.006-0.600 mg/mL was used as a control.

Chelating activity on Fe2+

Chelating activity of the examined extracts on

Fe2+ was measured according to the method of

Decker and Welch [16]. Plant extracts were dissolved

in ethanol, in an appropriate manner to obtain a

series of dilutions (10 to 40 mg plant material/mL).

Table 1. Retention time, MRM transition, collision energy (CE), and fragmentor voltage of 18 phenolic compounds

Phenolic acid Time, min MRM (m/z) CE / eV Fragmentor voltage, V

Quinic acid 0.612 191>127; 109 12;12 99

Gallic acid 1.399 169>125 8 103

Protocatechuic acid 1.787 153>109 8 67

p-Hydroxybenzoic acid 4.787 137>93 8 82

Chlorogenic acid 6.612 353>191 20 113

Vanillic acid 7.135 167>123; 152 4;8 93

Caffeic acid 7.615 179>107; 135 16;8 73

Syringic acid 9.873 197>106; 121 16;8 73

Ferulic acid 14.254 193>134; 149; 178 4;4;4 119

Sinapic acid 14.569 223>208; 164 4;0 118

Rosmarinic acid 14.569 359>161;179 40;56 170

Flavonoids

Catechin 1.070 289>245; 203; 137 8;12;16 134

Epicatechin 1.830 289>109 20 134

Rutin 3.192 609>179; 300 40;36 118

Myricetin 3.528 317>109; 137; 151 32;28;20 144

Quercetin 4.578 301>121; 151; 179 28;16;12 135

Naringenin 4.761 271>151; 119 8;20 93

Kaempferol 6.246 285>151; 93; 117 16;36;40 155


42

The IC50 value (mg/mL) was defined as the concen-

tration of extract that chelates 50% of Fe2+ under the

experimental conditions. It was obtained by interpol-

ation from linear regression analysis. EDTA in the

concentration range 0.006-0.600 mg/mL was used as

a control.

Reducing power

The reducing power of the examined extracts, in

series of different concentrations (0.100 to 15 mg

plant material/mL), was determined according to the

method of Oyaizu [17]. This method is based on the

reduction of Fe3+ to Fe2+ and measuring absorbance

(at 700 nm) of the Perl’s Prussian blue complex. The

IC50 value (mg/mL) was defined as an effective con-

centration of extract at which the absorbance of react-

ion mixture reaches 0.5 for reducing power. It was

obtained by interpolation from linear regression ana-

lysis. BHT in the concentration range 0.006-0.600

mg/mL was used as a control.

β-Carotene bleaching method

Antioxidant activity (AA) of the series of exam-

ined extracts (1 to 40 mg plant material/mL) was det-

ermined using β-carotene bleaching method, an in

vitro assay which measures oxidative loss of β-caro-

tene in β-carotene/linoleic acid emulsion. Thermal

autoxidation at 50 C was performed for 2 h. The deg-

radation rate of β-carotene was calculated according

to first-order kinetics and the AA was expressed as

percent of inhibition relative to the control [18]. The

IC50 value (mg/mL) was defined as the concentration

at which AA was 50% under the experimental con-

ditions, and it was obtained by interpolation from

linear regression analysis. BHT in the concentration

range 0.2-2 mg/mL was used as a control.

Antihemolytic activity

Antihemolytic activity of the extracts was det-

ermined using the method of Ko et al. [19], which was

further optimized for DBLF extracts as previously

described by Šarić et al. [20]. A solution of hydrogen

peroxyde (0.0625 vol.%) in phosphate-buffered saline

(PBS) solution was used instead of concentrated sol-

ution, and incubation time was set to 2 h instead of 4

h from the original method. The extent of the hemo-

lysis in every sample was calculated as: 100(Asample/

/Atotal hemolysis) = % of hemolysis. Ascorbic acid solution

in the concentration range 0.005-0.500 mg/mL was

used as a control. Since the percentage of hemolysis

was calculated for different sample concentrations (1

to 40 mg plant material/mL), these values were plot-

ted against sample concentrations and, using linear

regression, IC50 values (concentration of the inves-

tigated extract at which 50% of hemolysis inhibition is

achieved) of every investigated extract, as well as the

ascorbic acid solution, were calculated.

Schaal oven test

The Schaal oven test at 70 C was conducted to

evaluate the antioxidant effectiveness of ethanolic

extracts of DBLF in retarding the rancidity of com-

mercially available refined sunflower oil during 12

days of storage. For that purpose, 5 g of oil was

mixed with 20 mass% tested extracts and spread in 1

cm layer in glass containers. The test was carried out

in the dark and oxidative changes were monitored

gravimetrically. Experiments were also carried out

with synthetic antioxidant, BHT at 10 ppm level and

the control sample with no added antioxidants. The

extracts were evaporated to dryness and redissolved

in methanol. A control sample was prepared by using

the same amount of methanol used to dissolve the

antioxidant (BHT) and the extracts. In order to monitor

the kinetics of oxidation process [21], first derivatives

of samples weight gain were calculated using the

software SciDAVis 0.2.4. [22].

Statistical analysis

Apart from the extraction procedures and Schaal

oven test, which were done in duplicate, all analyses

were performed in triplicate, and the mean values

with the standard deviations are reported. Analysis of

variance and Duncan's multiple range tests were

used. Statistical data analysis software Statistica

(StatSoft, Inc. 2011) [23] was used for analysis. P

values < 0.05 were regarded as significant.


Referring to Kim et al. [24] who tested different

solvents in order to find the most effective way for the

extraction of rutin from buckwheat, the use of aque-

ous ethanol and acetone (both 50 vol.%) as extraction

solvents produced the highest yields of rutin. Kim et

al. [24] also found that the extraction temperatures in

the range 60–80 C and the extraction time from 0.5 to

1 h were optimal for achieving a high recovery yield of

rutin. On the other hand, Hinneburg and Neubert [12]

suggested that an extract with good antioxidant act-

ivity, a high content of phenolics, and a low content of

the phototoxic fagopyrin can be yielded by agitated

maceration with 30% ethanol at 60 C for 2 h.

Due to their low toxicity and supposed effici-

ency, water and ethanol/water mixtures were chosen

for the extraction of phenolic compounds from DBLF

in our experiment. Instead of prolonged heating which

is energy consuming and may result in destruction of


43

thermally labile constituents of interest, the effect of

short term heating, just to bring the mixture to boil and

24 h maceration afterwards on the extraction effici-

ency was examined. The following extraction yields

were obtained: water extraction = 20.34%; boiling

water extraction = 19.72%; ethanol/water (50:50)

extraction = 21.17%; boiling ethanol/water (50:50)

extraction = 21.28%; ethanol/water (80:20) = 21.28%

; boiling ethanol/water (80:20) extraction = 19.70%.

Total flavonoid content of plant extracts and HPLC

identification and quantification of phenolic

compounds

Besides the total flavonoid content determin-

ation, chemical characterization of the extracts inc-

luded identification of phenolic compounds by LC-

-MS/MS and quantification of identified compounds by

HPLC/DAD. Eighteen secondary biomolecules were

included into the identification (LC-MS/MS) method,

based on the availability of reference standards in the

laboratory. Method development started with the sel-

ection of precursor ions, which was done in MS2Scan

mode. The ionisation predominantly resulted in the

formation of [M−H]–. To assure high yield of precursor

ions while simultaneously preventing in-source frag-

mentation, fragmentor voltage (V) was optimised for

each compound. For this purpose, a standard mixture

was analyzed in MS2SIM mode, using fragmentor

voltages from 60 to 200 V in 10 V increments. In order

to optimise collision energy, the standard mixture was

subsequently analysed in Product Ion Scan mode,

using [M−H]– as precursors, optimal fragmentor volt-

age, and collision cell voltages ranging from 0-50 V

(in 10 V increments). Identification of the phenolic

constituents of the extracts was based on the com-

parison of the retention times and ionization patterns

in MRM mode with those of the external standards

(Table 1). Out of 18 analyzed compounds, the pre-

sence of 7 was confirmed in the extracts. After the

identification, phenolic compounds were quantified by

HPLC/DAD method, which was previously optimized

and validated for the determination of phenolic com-

pounds in crude medicinal plant extracts [14]. Content

of phenolic compounds in each extract is presented in

Table 2. Referring to the obtained results, the most

abundant phenolic compound in DBLF extracts was

rutin with the exception of Extract A where chloro-

genic acid was present in the highest concentration

(Table 2). The highest yield of rutin was obtained by

using boiling ethanol/water (80:20, V/V) as extracting

solvent, while the highest yields of chlorogenic acid

were obtained by water extraction. Ethanol/water

(50:50) mixture was shown to be the most efficient for

the extraction of investigated phenolic acids. Accord-

ing to Hinneburg and Neubert [12], quercetin origin-

ates as a product of rutin degradation by flavonol-3-

-O-β-heterodisaccharidase in buckwheat herb or in an

extract. Referring to the obtained results, quercetin

was present in all extracts at a concentration, which

was not highly influenced by the difference in applied

extraction procedures.

Our results are in accordance with the findings

of Hinneburg and Neubert [12], who reported that the

main phenolics of DBLF are rutin, chlorgenic acid,

and hyperoside and that rutin content in buckwheat

herb can be up to 8%. Besides rutin and quercetin,

the other flavonoids like vitexin, isovitexin, orientin

and isoorientin, which could exhibit 4-40% of total

antioxidant activity, were also reported to be present

in green parts of buckwheat [25]. However, due to the

lack of external standards, those compounds could

not have been quantified by applied HPLC method.

Instead, total flavonoid content of each extract was

estimated (Table 3). Due to the difference in applied

methods, these results differ from the results obtained

by HPLC. However, strong positive correlation

between HPLC-determined rutin and total flavonoid

contents was observed (r = 0.931, P < 0.05), indi-

cating that rutin is by far the most abundant flavonoid

in the extracts.

Table 2. Content of phenolic compounds (mg/g DBLF) in extracts: A-water; B-boiling water; C-ethanol/water (50:50); D-boiling ethanol/

/water (50:50); E-ethanol/water (80:20); F-boiling ethanol/water (80:20); values are means of three determinations ± standard deviation.

Values a,b,c,d, in each row with the same superscript are not significantly different (P < 0.05)

Compound Extract A Extract B Extract C Extract D Extract E Extract F

Gallic acid 0.052±0.009 a 0.037±0.004 a 0.874±0.065 c 0.864±0.03 c 0.779±0.021 b 0.766±0.071b

Protocatechuic acid 0.172±0.011a 0.175±0.016 a 0.266±0.021 c 0.236±0.019 b 0.209±0.014 b 0.228±0.048 b

Caffeic acid 0.214±0.013 a 0.240±0.027 a 1.084±0.101 c 0.953±0.091 b 0.988±0.077 b 0.966±0.068 b

Catehin 0.093±0.007 b 0.062±0.007 a 0.088±0.005 b 0.087±0.005 b 0.090±0.003 b 0.092±0.008 b

Chlorogenic acid 1.652±0.075 d 1.621±0.111 d 0.855±0.042b 0.950±0.036c 0.772±0.052 a 0.743±0.026 a

Rutin 0.251±0.013a 2.274±0.125b 40.35±0.693d 39.87±0.265d 34.93±0.213 c 49.94±0.623e

Quercetin 0.252±0.031 a 0.254±0.026 a 0.310±0.031 bc 0.306±0.034 bc 0.344±0.029bc 0.260±0.022 ab


44

Antioxidant activity (AOA) of the extracts

Due to the complexity of oxidative processes in

food matrices and cells and the versatile nature of

plant extracts a single method for the screening of

antioxidant potential is not recommended. Therefore,

the antioxidant effects of plant products must be

evaluated by combining at least two or more different

in vitro assays to obtain relevant data. With respect to

this, the antioxidant properties of the examined ext-

racts were tested through their ability to donate elec-

tons or H atoms, chelate/deactivate transition metal

ions and to inhibit the process of lipid oxidation.

Free radical scavenging activity of the extracts

was tested by applying the method towards long-lived

DPPH• while the electron-donating capacity was eval-

uated by measuring their reducing power. Referring to

the results (Table 3) expressed as IC50 values, the

extracts showed lower DPPH• scavenging activity and

reducing power than the reference compound BHT

(Table 4). In comparison with other investigated

samples, water extracts had significantly lower red-

ucing power and DPPH• scavenging activity than the

ethanol/water extracts, which were not significantly

different in both of the tests. Thermal treatment had a

significant positive influence only on water extraction.

Also, the correlation between reducing activity and

scavenging activity on DPPH• (Table 4) was highly

positive. Furthermore, reducing activity significantly

correlated with rutin and total flavonoid contents (Table

4). Similarly, strong positive correlation was observed

between scavenging activity on DPPH• and rutin and

total flavonoid contents (Table 4).

As the process of lipid oxidation occurs both in

the high-fat food products causing rancidity and in

living organisms resulting in cell damage, three dif-

ferent lipid model systems were used to measure the

antioxidant activity of the extracts: β-carotene bleach-

ing method, antihemolytic test and Schaal test.

β-Carotene bleaching test measures the loss of

the yellow colour of β-carotene due to its reaction with

radicals that are formed by linoleic acid oxidation in

an emulsion. Antioxidant activity of an extract in this

case refers to its ability to protect β-carotene from

oxidative damage. The test results (Table 3) indicated

the following order in antioxidant activities: boiling

ethanol/water (50:50) = ethanol/water (50:50) = boil-

ing ethanol/water (80:20) > ethanol/water (80:20) >

water extracts. The extracts with the highest activity

Table 3. Antioxidant activity and total flavonoid content of dried buckwheat leaf and flower (DBLF) extracts: A-water; B-boiling water; C-

ethanol/water (50:50); D-boiling ethanol/water (50:50); E-ethanol/water (80:20); F-boiling ethanol/water (80:20) and the control sub-

stances; values are means of three determinations ± standard deviation. Values in each column with the same superscript a,b,c,d are not

significantly different (P < 0.05)

Extract/control

substance

IC50 in mg plant material/mL or mg control substance/mL Total flavonoid content

g/100g plant material DPPH• scavenging

activity

Chelating activity

on Fe2+

Reducing

power

β-Carotene

bleaching method

Antihemolytic

activity

A 3.79±0.177c 272.8±31.1d 15.10±2.43d 5.04±0.083c 42.1±7.27d 0.22±0.026a

B 2.78±0.044b 64.8±0.961c 12.46±1.89c 4.76±0.400c 29.4±9.04c 0.72±0.035b

C 0.800±0.074a 34.4±0.559b 3.19±0.135b 1.88±0.142a 3.43±0.207b 2.57±0.031d

D 0.679±0.018a 25.7±1.01b 2.81±0.093b 1.87±0.054a 3.12±0.370b 4.16±0.032e

E 0.816±0.025a 39.8±0.240b 3.40±0.342b 3.82±0.197b 5.53±0.339b 2.42±0.021c

F 0.672±0.009a 30.6±0.580b 2.66±0.040b 1.78±0.003a 2.30±0.377b 4.92±0.026f

BHT 0.560±0.001a – 0.360±0.010a 1.78±0.060a – –

Ascorbic acid – – – – 0.18±0.010a –

EDTA – 0.039±0.001a –– – – –

Table 4. Correlation coefficients between IC50 values of antioxidant activity and phenolic compounds (total flavonoid and rutin) content

(P < 0.05)

Parameter Chelating activity on

Fe2+

Antihemolytic

activity

β-Carotene

bleaching method

Reducing

power

DPPH• scavenging

activity

Rutin content Not significant - 0.963 - 0.924 -0.9587 -0.974

Total flavonoid content Not significant - 0.874 - 0.905 -0.871 -0.879

DPPH• scavenging activity 0.871 0.999 0.867 0.996 -

Reducing power 0.823 0.995 0.880 - -

β-carotene bleaching method Not significant 0.881 - - -

Antihemolytic activity 0.871 - - - -


45

were as efficient as BHT. Significant correlation was

found between these results and rutin content and

total flavonoid content (Table 4). Also, these results

significantly correlated with scavenging activity on

DPPH• and the results of reducing power (Table 4).

Antihemolytic activity assay shows the ability of

the tested extracts to inhibit hydrogen-peroxide ind-

uced oxidation of the lipids in the phospholipid bilayer

of erythrocyte membranes and has the advantage of

using biological system instead of simpler membrane

models like phosphatidylcholine liposomes [26].

Although much higher than the IC50 of ascorbic acid,

calculated IC50 values of the investigated extracts

(Table 3) significantly correlated with the total flavo-

noid content and even better with the rutin content

(Table 4) of the samples, which indicates that flavo-

noids play a crucial role in preventing oxidative dam-

age to erythrocyte cell membranes in vitro. Also, the

highly significant correlation of antihemolytic activity

with scavenging activity on DPPH• and reducing

power (Table 4) of the extracts implies the major

mechanisms of antioxidant protection under the given

experimental conditions.

In the third model system, which could be rel-

evant to lipid oxidation processes in high-fat food pro-

ducts, sunflower oil was used as a lipid substrate. The

courses of oxidation were followed gravimetrically

and demonstrated by using derivative plots (Figure 1).

The inflection points obtained in this manner corres-

ponded to the maximum change in weight gain over

time and are convenient for the induction period det-

ermination. As it can be seen from the inflection

points in Figure 1, the oxidation process of the sun-

flower oil (control) began on the fifth day. With the

addition of DBLF ethanolic extracts, the induction

period in vegetable oil was prolonged. After BHT,

which showed the highest efficiency in delaying oxid-

ative changes (induction period of 8 days), boiling

ethanol/water (80:20) extract showed the highest acti-

vity prolonging the induction period of sunflower oil for

7.5 days. The ethanol/water (80:20) extract was cap-

able of prolonging the beginning of the oxidation

process until 7th day. However, boiling ethanol/water

(50:50) and ethanol/water (50:50) extracts were able

to postpone the oxidation process only for 5 and 5.5

days, respectively. Water extracts were not tested as

they had been previously demonstrated to be less

efficient by other tests (Table 3).

The ability of an antioxidant to chelate/deact-

ivate transition metal ions, which can catalyze hydro-

peroxide decomposition and Fenton-type reactions, is

considered an important mechanism of AOA, and

therefore the chelating activity test was used to eval-

uate the chelating efficiency of investigated extracts.

Obtained IC50 values indicate that investigated ext-

racts possess significant chelating activities and may

be able to play a protective role against oxidative

damage by sequestering Fe2+, but they showed much

lower chelating activity than the reference compound

EDTA. The water extracts were shown to be less effi-

cient than the ethanol/water extracts, which were not

statistically different (Table 3). The correlation

between Fe2+ chelating activity rutin and total flavo-

noid content was positive, but not statistically signific-

ant, which indicates that the other compounds pre-

sent in the extracts, along with flavonoids contribute

to the chelating activity. Chlorogenic acid was rep-

orted to be a powerful chelator of iron ions, capable of

forming iron/chlorogenic acid complex, which was not

capable of generating the hydroxyl radicals in Fenton

model system [27]. However, the correlation between

chlorogenic acid content and chelating activity of the

Figure 1. Antioxidant effectiveness of dried buckwheat leaf and flower (DBLF) ethanolic extracts tested by Schaal oven test. Changes in

the mass per day (%/day) of sunflower oil samples supplemented with: A–boiling ethanol/water (80:20) DBLF extract; B–ethanol/water

(80:20) DBLF extract; C–boiling ethanol/water (50:50) DBLF extract; D–ethanol/water (50:50) DBLF extract; E–control;

F–0.01% BHT during accelerated oxidation.


46

extracts was not established. On the contrary, the

results of chelating activity significantly correlated

with those of scavenging activity on DPPH• and red-

ucing power (Table 4). Unlike the poor correlation

with β-carotene bleaching test, the results of anti-

hemolytic activity assay significantly correlated with

those of chelating activity on Fe2+ (Table 4). Contrary

to the β-carotene bleaching test, the antihemolytic

activity assay represents an iron ion-dependent lipid

peroxidation system, where a potential of an antioxid-

ant to act as a chelator becomes important [28].

CONCLUSION

DBLF was shown to be a rich source of anti-

oxidants. Extraction with ethanol/water mixture (80:20

V/V) for 24 h at room temperature, after the mixture

was just brought to boil was demonstrated to be an

efficient and cheap way for the extraction of antioxid-

ants.

Flavonoids, primarily rutin, were shown to be the

most responsible for the antioxidant activity in all

investigated lipid model systems, acting as free rad-

ical scavengers, electron-donating substances and

iron ion chelators. In the β-carotene bleaching test,

the extracts with the highest activity were as efficient

as BHT. Furthermore, the extracts were able to pro-

tect erythrocyte cell membranes from oxidative dam-

age and to prolong the induction period of sunflower

oil for 3.5 days under the accelerated oxidation con-

ditions.

Finally, the DBLF extracts obtained by this rel-

atively cheap and environmentally friendly extraction

procedure show a very good potential for incorpor-

ation in certain fat-containing food products as a rich

source of natural antioxidants, which, beside their

already proven health benefits, would provide a sub-

stantial protection against oxidative-induced rancidity.

However, in order to fully optimize the extraction

method, further research, which would include testing

of ratio of liquid to raw material, extraction time, the

ultrasonic power and radiation time on the extraction

efficiency needs to be done.

Acknowledgment

This work is a part of the National Project (TR-

-31029) financially supported by the Ministry of

Education, Science and Technological Development,

Republic of Serbia.

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Y.-B. Park, M.-S. Choi, Food Chem. 119 (2010) 235-240

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Pharm. Bull. 36 (1988) 2090-2097

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[17] M. Oyaizu, Jpn. J. Nutr. 44 (1986) 307-315

[18] M.S. Al-Saikhan, L.R. Howard, J.C.Jr. Miller, J. Food Sci.

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47

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ALEKSANDRA MIŠAN1

BOJANA ŠARIĆ1

IVAN MILOVANOVIĆ1

PAVLE JOVANOV1

IVANA SEDEJ1

VANJA TADIĆ2

ANAMARIJA MANDIĆ1

MARIJANA SAKAČ1

1Institut za prehrambene tehnologije u

Novom Sadu (FINS), Univerzitet u

Novom Sadu, Bulevar cara Lazara 1,

21000 Novi Sad, Srbija 2Institut za proučavanje lekovitog bilja

„Dr Josif Pančić“, Tadeuša Košćuška 1,

11 000 Beograd, Srbija

NAUČNI RAD

FENOLNI PROFIL I ANTIOKSIDANTNA SVOJSTVA EKSTRAKATA OSUŠENIH LISTOVA I CVETOVA HELJDE

Formulacije na bazi osušenih listova i cvetova heljde (DBLF) su dokazano efikasne u

tretiranju bolesti krvnih sudova, a što se dovodi u vezu sa visokim sadržajem rutina

(2-10%). U cilju pronalaženja ekonomičnog načina ekstrakcije antioksidanata testirani su

uticaj odnosa etanola i vode, kao i temperature na efikasnost ekstrakcije fenolnih jedinjenja

i mehanizme antioksidantnog delovanja ekstrakata. Ekstrakcija smešom etanol/voda

(zapreminski odnos 80:20) tokom 24 h na sobnoj temperaturi, nakon što je smeša zagre-

jana do ključanja, pokazala se kao efikasan i ekonomičan način za dobijanje visokog sadr-

žaja rutina (49,94±0,623 mg/g DBLF). Najzastupljenija fenolna jedinjenja u DBLF ekstrak-

tima bila su rutin i hlorogenska kiselina. Pokazano je da su flavonoidi, a posebno rutin,

najodgovorniji za antioksidantnu aktivnost ekstrakata u lipidnim model sistemima, delujući

kao “hvatači” slobodnih radikala, elektron-donorske supstance i helatori jona gvožđa. U

testu ispitivanja antioksidantne aktivnosti β-karoten metodom najpotentniji ekstrakti bili su

jednako efikasni kao i BHT (butilovani hidroksitoluen). Rezultati antihemolitičkog i Schaal

testa ukazali su na izuzetnu sposobnost ekstrakata da inhibiraju oksidativnu razgradnju

eritrocita i produže period do početka oksidativnih promena u suncokretovom ulju.

Ključne reči: osušeni listovi i cvetovi heljde; rutin; antioksidantna aktivnost; oksi-

dacija lipida; ekstrakcija.






49

YAJING ZHANG1,3

YU ZHANG1

FU DING1,3

KANGJUN WANG1,3

XIAOLEI WANG2

BAOJIN REN1

JING WU1

1College of Chemical Engineering,

Shenyang University of Chemical

Technology, Shenyang, China 2School of Science, Shenyang

University of Technology,

Shenyang, China 3Liaoning Co-innovation Center of

Fine Chemical Industry, Shenyang,

China

SCIENTIFIC PAPER

UDC 66.094.25:546.264–31:544.47

https://doi.org/10.2298/CICEQ150711005Z

SYNTHESIS OF DME BY CO2 HYDROGEN-ATION OVER La2O3-MODIFIED CuO–ZnO–ZrO2/HZSM-5 CATALYSTS

Article Highlights

• La-modified CuO-ZnO-ZrO2/HZSM-5 catalysts were prepared by an oxalate co-precipit-

ation method

• La-modified CuO-ZnO-ZrO2/HZSM-5 catalysts show higher catalytic performances

• The catalytic performances of the catalysts are strongly dependent on the La content

Abstract

A series of La2O3-modified CuO-ZnO-ZrO2/HZSM-5 catalysts were prepared by

an oxalate co-precipitation method. The catalysts were fully characterized by

X-ray diffraction (XRD), N2 adsorption-desorption, hydrogen temperature pro-

grammed reduction (H2-TPR), ammonia temperature programmed desorption

(NH3-TPD), and X-ray photoelectron spectroscopy (XPS) techniques. The effect

of the La2O3 content on the structure and performance of the catalysts was

thoroughly investigated. The catalysts were evaluated for the direct synthesis of

dimethyl ether (DME) from CO2 hydrogenation. The results displayed that La2O3

addition enhanced catalytic performance, and the maximal CO2 conversion

(34.3%) and DME selectivity (57.3%) were obtained over the catalyst with 1%

La2O3, which due to the smaller size of Cu species and a larger ratio of Cu+/Cu.

Keywords: CO2 hydrogenation; dimethyl ether; La2O3 promoter.

Carbon dioxide emission has caused irrever-

sible climate changes due to its greenhouse effect [1].

CO2 capture and storage is one of efficient method for

reducing the CO2 emissions. However, this techno-

logy requires high cost [2]. Chemical recycling of CO2

is another main option, which is presented as an eco-

nomically attractive and sustainable method. CO2 can

be converted into methanol, dimethyl ether (DME),

methane and syngas (CO+H2) [3], carboxylic acids,

etc. Among them, synthesis of DME has been paid

special attention since DME can be applied as a clean

fuel, coolant, propellant, and is an important chemical

intermediate [4]. The chemical reactions occurring in

direct conversion of CO2 to DME can be described by

the following equations [5]:

Correspondence: F Ding, K. Wang, College of Chemical Eng-

ineering, Shenyang University of Chemical Technology, Shen-

yang 110142, PR China. E-mail: F. Ding, [email protected];

K. Wang, [email protected] Paper received: 11 July, 2015 Paper revised: 8 January, 2016 Paper accepted: 19 February, 2016

H2 2 3 2CO 3H CH OH H O, 49.4kJ/ mol (1)

H3 3 3 22CH OH CH OCH H O, 23.4 kJ / mol (2)

H2 2 2CO H CO H O, 41.2 kJ/ mol (3)

For the methanol synthesis (reaction (1)), CuO–

–ZnO–ZrO2 catalytic system is considered as more

favorable than the traditional CuO–ZnO–Al2O3, because

the interaction of Cu metal particles with ZnO and

ZrO2 lead to the stabilization of a “mix” of Cu and Cuδ+

(not Cu2+, Cu1+ and Cu) [6]. For the methanol dehyd-

ration process (reaction (3)), solid-acid catalysts

(HZSM-5, γ-Al2O3 and sulfated zirconia) are employed

[7]. HZSM-5 is widely applied for methanol dehyd-

ration because: 1) it has very high catalytic activity at

the optimum reaction temperature; 2) it is more res-

istant toward poisoning of acid sites by the water due

to more hydrophobic character; 3) predominance of

Brønsted-type acidity, which can promote the DME

yield [8]. However, H-ZSM-5 presents the disadvant-

ages of narrow pore size and strong acid sites, which

Y. ZHANG et al.: SYNTHESIS OF DME BY CO2 HYDROGENATION… Chem. Ind. Chem. Eng. Q. 23 (1) 4956 (2017)

50

will limit reactant molecules diffusion and lead to

formation of secondary products, respectively. Witoon

et al. developed sulfated zirconia catalyst with 20%

sulfur-loaded on ZrO2 as methanol dehydration cat-

alyst, the yield of methanol and DME over CuO–ZnO–

–ZrO2/20S-ZrO2 both higher than those over CuO–

-ZnO–ZrO2/H-ZSM-5, but the stability of the former

catalyst is a little poor [9].

Up to now, the conversion of CO2 and the sel-

ectivity of DME are still not high. To enhance the cat-

alytic activity, efforts have been addressed by adding

promoters for methanol synthesis catalyst, besides

developing new preparation methods [10-14]. Our

previous results showed that V and Mn oxides mod-

ified CuO–ZnO–ZrO2/HZSM-5 exhibited higher cat-

alytic performance [15-16]. La2O3, as a rare earth

metal oxide, is considered as owing some basic char-

acter [17], and it can promote many metal oxide cat-

alytic reactions. Guo et al. [18] investigated catalysis

performance of La doping Cu/ZrO2 for CO2 hydro-

genation to methanol, and they found the amount of

basic site increases with La loading and the presence

of La enhance the selectivity of methanol. Gao et al.

reported appropriate amount of La can decrease the

crystallite size of CuO and enhance the dispersion of

Cu [19]. Sun’s group has investigated the effect of La

on the performance of Cu/Zn/Al catalysts via hydro-

talcite-like precursors for CO2 hydrogenation to meth-

anol, and the results showed La addition not only led

to higher BET specific surface area and Cu disper-

sion, but also increased the total number of basic

sites and proportion of strongly basic sites [20].

Furthermore, Sun’s group also reported that addition

of Zr to La-Cu-Zn-O with perovskite structure catalyst,

for synthesis of methanol from CO2 hydrogenation,

will lead to smaller particles, lower reduction tempe-

rature, higher Cu dispersion, larger amount of hydro-

gen desorption at low temperature as well as higher

concentration of basic sites are obtained [21]. How-

ever, the introduction of La2O3 into the CuO–ZnO–

-ZrO2/HZSM-5 catalyst for CO2 to DME has not been

reported.

Herein, La2O3-modified CuO-ZnO-ZrO2/HZSM-5

catalysts with a different La2O3 content were pre-

pared, aiming at investigating the effect of La2O3 mod-

ification on the structure and performance the cat-

alysts.

EXPERIMENTAL

Catalyst preparation

The La2O3 modified CuO-ZnO-ZrO2/HZSM-5

(CuO:ZnO:ZrO2 mass ratio: 5:4:0.2) catalysts, abbre-

viated as CZZLxH where x stands for theoretical

La2O3/CZZ wt.%, were prepared by an oxalate co-pre-

cipitation method (CZZLx/HZSM-5 mass ratio: 2:1),

HZSM-5 (SiO2/Al2O3 mole ratio: 50) was purchased

from Catalyst Plant of Nankai University (China). The

preparation method is same as described in other

work [15]. In brief, first, metal nitrates were dissolved

into a certain amount of ethanol (denoted as solution

A); H2C2O42H2O (200 mol.% of metal nitrate) was

also dissolved into ethanol (solution B). Then, sol-

utions A and B were slowly dropped into a beaker

containing a HZSM-5 suspension in ethanol suspen-

sion kept under stirring at 333 K. The suspension was

sealed and aged for 2 h and then the ethanol was

evaporated at 353 K to get a precipitate. Finally, the

precipitate was dried at 393 K for 12 h and calcined in

air at 673 K for 4 h.

Catalyst testing

Catalytic performance was evaluated in a con-

tinuous-flow fixed-bed reactor made of stainless steel

with inside diameter of 0.01 m. First, the catalyst was

reduced with 10% H2/N2 at 573 K for 3 h under atmo-

spheric pressure. Then it was cooled to 653 K and

reactant gas flow was introduced, raising the pressure

to 3.0 MPa, the reaction temperature was 543 K. The

exit line was heated to prevent condensation. The

products were analyzed on line with a gas chroma-

tograph (SP2100A) equipped with both a TCD (for CO

and CO2, GDX-101 connected with Porapak T col-

umn) and a FID (for CH4, CH3OH and CH3OCH3,

Porapak Q column). Conversion and selectivity

values were calculated by internal standard method

[22]. XCO2, Spi and YDME represent the conversion of

CO2, the selectivity of the product (DME, MeOH and

CO) and the yield of DME, respectively. Each expe-

rimental data was corresponds to an average of three

independent measurements, with error of ±2%.

2

2,in 2,out

2,in 2,outCO

2,in

2,in

CO CON N

CON

X (4)

iPi

2,out1 CO

PS (5)

where Pi stands for the concentration of a specific i

product (DME, MeOH orCO):

2DME DME COY S X (6)

Catalyst characterization

XRD measurements were performed on a Rig-

aku D/max 2500pc X-ray diffractometer with Cu-Kα


51

radiation ( = 1.54156 Å) at a scan rate of 4 min-1 at

50 kV and 250 mA. For the reduced catalysts were

first reduced with 10% H2/N2 at 573 K for 3 h, and

cooled to room temperature under the N2 flow, and

then keep it into a bottle full of N2 and send it to XRD

chamber immediately. The crystallite size was calcul-

ated using the Scherrer equation. BET surface areas

were measured by N2 adsorption at 77 K using a

Quantachrome Autosorb 1-C. Before measurements,

samples were degassed under vacuum at 573 K for 4

h. H2-TPR was carried out in 10% H2/Ar flowing at 50

mL min-1, using a ramp rate of 10 K min-1 to 773 K.

NH3-TPD was conducted on a Chemisorb from 373 to

873 K. XPS measurements were performed on an

ESCALAB-250. The catalysts were first reduced with

10% H2/N2 at 573 K for 3 h, and cooled to room tem-

perature under the N2 flow, then put into the chamber

of X-ray photoelectron spectrometer immediately for

measurement. The exact composition of the surface

of the catalyst was determined by XPS. Acidity on the

catalyst was measured on a NICOLET 500 FT-IR

through pyridine adsorption. The sample was pre-

pared to the load slice, which was subject to purific-

ation under vacuum pressure 0.0133 Pa at 673 K,

after cooling down to the room temperature, pyridine

was adsorbed for 0.5 h. Desorption was carried out by

programming temperature to 423 K. The infrared

spectrum was generated within the scope of 1400–

–1700 cm–1 at room temperature.


Catalytic performance of catalysts

The catalytic performances of CZZLxH catalysts

with varying La2O3 content are summarized in Table

1. The major product was DME, and the side products

were methanol, CO and trace hydrocarbons. The

amount of hydrocarbons was less than 1% and there-

fore they were neglected. The CO2 conversions and

DME selectivities over the La2O3-modified catalysts

are higher than those over unmodified one (CZZL0H),

indicating that La2O3-modification can efficiently

enhance the catalytic performances. The CZZL1H

exhibited the maximum CO2 conversion and DME sel-

ectivity of 34.3 and 57.3%, respectively. CO2 conver-

sion over CZZL1H increases 18.6%, compared with

CZZL0H. The influence of the catalyst composition on

the performance does not seem distinct, which is due

to the thermodynamic regime of reaction test because

of combination of high temperature and long contact

time. If the temperature is decreased to 533 K, the

differences of conversion and DME selectivity between

CZZL1H and CZZL0H become much larger (Table 2).

It is also noted that the selectivity of CO is high (about

30%), which may due to the higher reaction tempe-

rature of 543 K. CO was produced by reverse water-

gas reaction, the reaction has endothermic character,

as shown in reaction (2). Meanwhile, compared to

methanol synthesis, the RWGS reaction has a higher

apparent activation energy [23], indicating that the

increase in CO production is faster than that of meth-

anol with higher reaction temperature. The selectivity

of CO over CZZL1H exhibited the minimal value, indi-

cating that a certain amount of La2O3 can inhibit the

RWGS reaction. However, as Gao et al. pointed out,

further work needs to be carried out to investigate

how La2O3 changes the RWGS reaction [19].

In order to further increase the DME yield, the

catalytic performance of the CZZL1H catalyst was

also investigated at lower gas hourly space velocity.

As shown in Table 2, when the GHSV was as low as

1800 h-1, the CO2 conversion and DME selectivity

increase to 36.4 and 58.2%, respectively. The inc-

rease of the conversion can be attributed to the longer

contact time of CO2 and H2 over the catalyst at lower

GHSV. Similarly, methanol dehydration can proceed

to a higher degree with increasing contact time, lead-

ing to higher DME selectivity. In addition, a contrast

experiment was carried out at a reaction pressure of 5

MPa while keeping other conditions constant. DME

selectivity increased by approximately 7.3% and CO

selectivity decreased by 24.6% compared with those

obtained at 3 MPa. This result suggests that an inc-

rease of reaction pressure can improve the catalytic

performance.

The structure of catalyst

Figure 1A shows the XRD patterns of CZZLxH

catalysts. The peaks appearing at 35.5, 38.7, 48.7,

Table 1. Catalytic performances of the catalysts; reaction conditions: T = 543 K; p = 3.0 MPa; CO2:H2 = 1:3; GHSV = 4200·h-1

Catalyst Conversion of CO2, % Selectivity, %

DME Yield, % DME Productivity, g g cat-1 h-1 DME CH3OH CO

CZZL0H 28.9 55.1 12.9 32.0 15.9 0.171

CZZL0.5H 30.4 55.2 12.4 31.7 16.8 0.181

CZZL1H 34.3 57.3 13.3 29.4 19.6 0.212

CZZL2H 31.1 53.6 12.7 33.7 16.7 0.180


52

and 66.2 can be ascribed to CuO phase (tenorite,

JCPDS 48-1548), the peaks at 31.7, 34.3, 36.6, 56.5,

62.8 and 67.8 can be indexed to ZnO phase (JCPDS

65-3411), and the peaks appearing in the 2 range of

21-25 belong to HZSM-5 (JCPDS 44-0003). No

peaks belonging to ZrO2 or La2O3 are observed, indi-

cating that ZrO2 and La2O3 are amorphous or well

dispersed in catalyst body. The intensities of the

peaks assigned to CuO and ZnO weakened and the

widths of the peaks broadened gradually with inc-

reasing La2O3 content from 0 to 1%, then they

became sharper and narrower again when La2O3

content was 2%. This result indicates that the addition

of a proper amount of La2O3 can enhance the dis-

persion of CuO and ZnO [24]. This changing trend

can be well reflected in the changing trends of Cu and

CuO grain size, as shown in Table 3.

Table 3. Physicochemical properties of the catalysts; diffraction

peak at 2θ 38.7 for CuO and 43.3 for Cu

Catalyst SBET / m2 g–1 DXRD / nm

CuO Cu

CZZL0H 135.4 11.3 13.5

CZZL0.5H 136.7 9.6 11.9

CZZL1H 138.2 8.4 10.5

CZZL2H 135.8 9.3 11.4

The XRD patterns of the reduced catalysts are

shown in Figure 1B. The diffraction peaks at 2θ

values of 43.3, 50.4 and 74.1 can be indexed to the

crystal planes of (111), (200) and (220) of metallic

copper phase, respectively (JCPDS 04-0836). No dif-

fraction peaks belonging to the CuO phase could be

detected, suggesting all CuO species had been red-

uced to copper. The intensities of the peaks assigned

to Cu and ZnO changed in a similar trend as those of

CuO and ZnO (Figure 1A).

The specific surface areas and the calculated

crystallite sizes of the catalysts using Scherrer’s

equation are listed in Table 3. The SBET increased

from 135.4 of CZZL0H to 138.2 m2 g-1 of CZZL1H, and

then decreased to 135.8 m2 g-1 for CZZL2H. The

changing trend of the DCu is opposite to the trend of

the SBET, a minimum of 10.5 nm is obtained over

CZZL1H. Although the SBET change is not as signific-

ant as DCu, the trend is in accordance with XRD

results. Combining the results in both Tables 1 and 3,

it can be observed that CZZL1H shows the smallest

Cu crystallite size and the best catalytic performance,

which indicates the activity is closely related to the

crystallite size of Cu. Guo et al. also reported the cat-

alytic performance of Cu-TiO2-ZrO2 related to the

crystallite size of CuO [25].

Table 2. Catalytic performances of the catalysts at different reaction conditions; reaction conditions: CO2:H2 = 1:3

Catalyst T

K

Conversion of CO2

%

p

MPa

GHSV

h-1

Selectivity, % DME Yield

%

DME Productivity

g g cat-1 h-1 DME CH3OH CO

CZZL1H 543 34.3 3 4200 57.3 13.3 29.4 19.6 0.212

CZZL1H 543 36.4 3 1800 58.2 14.2 27.6 21.1 0.098

CZZL1H 543 38.5 5 1800 62.5 16.7 20.8 24.0 0.111

CZZL0H 533 24.2 3 4200 57.1 13.7 29.2 13.8 0.148

CZZL1H 533 30.4 3 4200 61.8 14.2 24.0 18.8 0.203

20 30 40 50 60 70 80

d

b

c

Inte

nsi

ties

(a. u

.)

2Theta (deg)

a

CuO

ZnO

HZSM-5

A

20 40 60 80

B

**

*

Cu*

ZnOHZSM-5

c

b

d

a

Inte

nsi

ties

(a.

u.)

2 Theta (deg)

Figure 1. A) XRD patterns of La2O3-modified CuO–ZnO–ZrO2/HZSM-5 and B) reduced catalysts: a) CZZL0H;

b) CZZL0.5H; c) CZZL1H; d) CZZL2H.


53

The reducibility of catalyst

H2-TPR results are depicted in Figure 2. It is

obvious that La2O3 modification displays a significant

effect on the interaction between the metal and the

carriers. Each catalyst exhibits a broad peak with

peak maximum in 423-563 K, corresponding to the

reduction of CuO to Cu [26]. For the catalysts with

increasing La2O3 content from 0 to 0.5 or 1%, the

peaks maxima shift to higher temperature from 529 to

554 or 556 K while no shape changes are observed,

which indicates the interaction between ZnO and CuO

or La2O3 and CuO becomes stronger [19,23,27]. This

result seems conflicting with XRD result. According to

XRD result, the dispersion of CuO becomes better

and the crystalline size becomes gradually smaller

with La2O3 addition amount from 0 to 1%, suggesting

that the reducibility of CuO should become easier.

This fact implies La2O3 modification still plays another

role. For CZZL2H catalyst, it is evident that the H2-

-TPR change tendency becomes more different and

the peak maximum shifts towards the lower tempe-

rature, indicating the reducibility of CuO becomes

easier [28,29]. It is concluded that La2O3 modification

has at least two roles. On one hand, La modification

promotes the dispersion of CuO, leading to easier

reduction of CuO; on the other hand, La2O3 modific-

ation enhances the interaction between CuO and

other metal oxides resulting in a more difficult red-

uction process. The two effects compete with each

other and the reduction temperature is dependent on

which effect is predominant. Xiong et al. also found

that the La2O3 can increase the dispersion of Co/AC

catalysts, whereas the reduction temperature shifted

to a higher position due to stronger interaction [30].

350 400 450 500 550 600 650

a

b

c

Co

nsu

mp

tio

n o

f H

2 (

a.u

.)

Temperature (K)

d

Figure 2. H2-TPR profiles of catalysts: a) CZZL0H; b) CZZL0.5H;

c) CZZL1H; d) CZZL2H.

Surface acidity of catalyst

Figure 3 shows the NH3-TPD results obtained

for pure HZSM-5 and CZZLxH catalysts. The total

acidic amount, the strength and the fraction of various

acid sites are summarized (supporting info, available

from the authors upon request). On pure HZSM-5 pro-

file, two NH3 desorption peaks are observed, indicat-

ing the existence of at least two different acid

strengths. In general, the peak located in 393-523 K

and 573-773 K can be attributed to weak and strong

acid strengths, respectively [31]. But for all catalysts,

three NH3 desorption peaks, in the temperature

regions of 373-473 K, 473-573 K and 573-673 K are

observed, denoted as α, β and γ peak, which can be

assigned to weak, medium and strong acid strengths

of HZSM-5, respectively [32]. This result is similar

with that of the reported V-modified catalyst [15].

According to the study, the reason why the catalysts

showed another more NH3 desorption peak than pure

HZSM-5 is mainly ascribed to the fact that strong acid

strength on pure HZSM-5 are blocked and modified

by metal oxides and oxalic acid, respectively. Com-

pared with La2O3-free catalyst, peaks α and β on the

other curves shift a little to higher temperature with

the increasing La2O3 content, implying that the weak

and medium acid strengths become stronger; on the

contrary, the peak γ shifts a little towards lower tem-

perature, indicating strong acid strength becomes

weaker. It is also worth noting that the total acid

amount decreases with the increasing amount of

La2O3 from 0.5 to 1%, which can be explained by the

basic character of La2O3. Sugi et al. also observed

this effect and suggested that La2O3-modification

resulted in reducing the support acidity [33]. However,

continuous addition of La2O3 to 2% increases the total

400 500 600 700

427K

421K

597K

500 K

599K499 K

656K

461 K

e

d

c

b

a

Am

mo

nia

des

orp

tio

n/a

.u

Temperatue (K)

Temperature, K

Figure 3. NH3-TPD profiles of pure HZSM-5 and the catalysts:

a) HZSM-5; b) CZZL0H; c) CZZL0.5H; d) CZZL1H; e) CZZL2H.


54

acid amount again. Previously published studies pro-

posed that there are two reasons for the acidity

amount increase [34]. One is that the La3+ has some

Lewis acidic property originated from an empty f orb-

ital; the other reason is that Si-OH and Al-OH in the

zeolite framework are polarized by La3+, which result

in a stronger acidity. Therefore, the acidity of the

CZZLxH catalysts is dependent on the La2O3 content.

Moreover, La2O3 modification can also affect the

concentration and distribution of the three acid

strengths of HZSM-5 in the catalysts. With the inc-

reasing amount of La2O3, the concentration and fract-

ion of medium acid strength increase and those of

strong acid strength become smaller as compared to

those on CZZL0H, but the difference is too small

(supporting info). It is generally considered that the

strong acidic strength on HZSM-5 zeolite promotes

the generation of secondary products like hydrocar-

bons, resulting in low selectivity to DME [6]. However,

here the change of the acidity could not be a factor

accounting for the higher DME selectivity because the

methanol selectivity of the various catalysts was

almost the same, as shown in Table 1. This result

suggests that the acidity of the catalysts is strong

enough to efficiently convert the produced methanol

to DME. In addition, the acidity of Brønsted acid sites

and Lewis acid sites for CZZL0H and CZZL1H cat-

alysts were determined (supporting info, available

from the authors upon request). The results show the

acidity of both type acid sites of CZZL1H decreased

compared CZZL0H, thus it can be inferred that there

is no direct relationship between the selectivity of

DME and the change of acid type of the two catalysts.

So, the improvement of DME selectivity is due to the

decrease in CO selectivity resulted from the Cu-based

catalyst but not from the acid component of the

bifunctional catalysts. In addition, according to the lit-

erature, it could be speculated that the introduction of

La2O3 into catalyst will increase the surface basicity of

catalyst, which in turn promotes the adsorption of CO2

and sequence enhances the yield for methanol, finally

DME selectivity is increased after methanol dehyd-

ration [18].

Results of XPS investigations

The reduced CZZLxH catalysts characterized by

XPS, the binding energy of Cu2p3/2, Zn2p3/2, as well

as the surface compositions of the catalysts are sum-

marized (supporting info, available from the author

upon request). For all the reduced catalysts, binding

energies (BE) of Cu2p3/2 are located at about 932.3

eV, which are the characteristic peaks of reduced

Cu+/Cu species [35]. The binding energy shifted to

higher positions with increasing amounts of La2O3,

which indicated stronger interaction between CuO

and other metal oxides carriers [36]. For the purpose

of distinguishing Cu+ from Cu species, their kinetic

energies in the XAES Cu LMM line positions were

measured (Figure 4). The Cu LMM spectra show a

broad and asymmetrical peak, implying the coexist-

ence of Cu+ and Cu in the surface of the catalysts.

Two symmetrical peaks centered at near 916.6 and

918.7 eV can be obtained by deconvolution, which

are corresponding to Cu+ and Cu species, respect-

ively [37]. Additionally, Cu+/Cu can be calculated

based on the results. Volcanic shape change trends

of the Cu+/Cu versus La addition content are obs-

erved, the CZZL1H exhibited the maximum of 0.144,

which may consequently lead to higher activity due to

the stabilization of Cu+ favoring the hydrogenation of

CO2 [38]. The binding energies (BE) of Zn2p3/2 for all

catalysts are located at about 1021.8 eV, which are

close to the characteristic peaks of ZnO species [39].

Compared to the nominal surface compositions of the

catalysts, it can be seen that the actual surface com-

position Cu/Zn decreased significantly, implying

enrichment of Zn. Similarly, La2O3 content is much

higher on the surface. It is worth noting that is La2O3

is not detectable in CZZL0.5H, possibly because the

content is too small to give enough signal.

915 918 920

d

c

b

a918.7916.6

Inte

nsi

ties

(a. u

.)

Kinetic Energy (eV)

Figure 4. Cu LMM XAES spectra of the reduced catalysts:

a) CZZL0H; b) CZZL0.5H; c) CZZL1H; d) CZZL2H.

CONCLUSIONS

La2O3 modification has great impact on the cat-

alytic performance of CuO-ZnO-ZrO2/HZSM-5 cat-

alysts for promoting direct CO2-to-DME. La2O3-mod-

ification can efficiently enhance the catalytic perform-

ance of CuO-ZnO-ZrO2/HZSM-5 catalysts. The

sample containing a nominal amount of 1% La2O3


55

gave the maximum CO2 conversion and DME sel-

ectivity of 34.3 and 57.3%, respectively, benefiting

from smaller Cu particles and a larger Cu+/Cu ratio.

When the La2O3 content is low (from 0.5 to 1%), it can

strengthen the interaction between CuO and other

metal oxides and inhibit the reduction of CuO species;

meanwhile, the total acid amount decreases slightly

but the medium strong acid concentration and strength

increase a little. However, excess of La2O3 content,

e.g., 2%, will lead to an opposite effect. In summary,

suitable La2O3 addition can improve the catalytic per-

formance of CuO-ZnO-ZrO2/HZSM-5 for one step

CO2 to DME transformation.

Acknowledgement

The authors thank National Nature Science

Foundation of China (51301114, 21201123,

21203125 and 61403263), SRF for ROCS, SEM (No.

[2010]1174), Natural Science Foundation of Liaoning

Province (2015020649), Liaoning Educational Depart-

ment Foundation (L2013161), LNET (LJQ2013044)

for financial support.

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56

YAJING ZHANG1,3

YU ZHANG1

FU DING1,3

KANGJUN WANG1,3

XIAOLEI WANG2

BAOJIN REN1

JING WU1

1College of Chemical Engineering,

Shenyang University of Chemical

Technology, Shenyang, China 2School of Science, Shenyang Univer-

sity of Technology, Shenyang, China 3Liaoning Co-innovation Center of Fine

Chemical Industry, Shenyang, China

NAUČNI RAD

SINTEZA DIMETIL ETRA HIDROGENIZACIJOM CO2 POMOĆU LA2O3-MODOFOKIVANIH CuO–ZnO–ZrO2/HZSM-5 KATALIZATORA

Oksalatnom ko-precipitacionom metodom je pripremljena serija katalizatora CuO-ZnO-

–ZrO2/HZSM-5 modifikovanih pomoću La2O3. Katalizatori su okarakterisani X-difrakcionom

metodom (XRD), N2 adsorpcijom-desorpcijom, termoprogramiranom redukcijom (H2-TPR),

termoprogramiranom desorpcijom (NH3-TPD) i fotoelektronskom spektroskopijom X-zraka

(XPS). Istražen je i uticaj sadržaja La2O3 na strukturu i performanse katalizatora. Kataliza-

tori su testirani u procesu CO2 hidrogenizacije i direktne sinteze dimetil etra (DME). Rezul-

tati pokazuju da dodatak La2O3 poboljšava performanse katalizatora, kao i da su maksi-

malna konverzija CO2 (34,3%) i DME selektivnost (57,3%) dobijeni upotrebom katalizatora

sa 1% La2O3, zbog manjih Cu čestica i veceg odnosa Cu+/Cu0.

Ključne reči: hidrogenizacija CO2, dimetil etar La2O3 promoter.






57

TATJANA

KALUĐEROVIĆ RADOIČIĆ1

NEVENKA

BOŠKOVIĆ-VRAGOLOVIĆ1

RADMILA GARIĆ-GRULOVIĆ2

MIHAL ĐURIŠ2

ŽELJKO GRBAVČIĆ1

1Faculty of Technology and

Metallurgy, University of Belgrade,

Belgrade, Serbia 2Intitute of Chemistry, Technology

and Metallurgy, Department for

Catalysis and Chemical

Engineering, University of

Belgrade, Belgrade, Serbia

SCIENTIFIC PAPER

UDC 66.021.1:621.798

https://doi.org/10.2298/CICEQ150506006K

FRICTION FACTOR FOR WATER FLOW THROUGH PACKED BEDS OF SPHERICAL AND NON-SPHERICAL PARTICLES

Article Highlights

• Experimental evaluation of pressure drop correlations in packed beds was conducted

• Pressure drop across beds of spherical and non-spherical particles was measured

• Spherical glass particles and quartz filtration sand were used as packing material

• Correlations in the form of Ergun equation gave the best results

• The coefficients in Ergun equation are system-specific

Abstract

The aim of this work was the experimental evaluation of different friction factor

correlations for water flow through packed beds of spherical and non-spherical

particles at ambient temperature. The experiments were performed by mea-

suring the pressure drop across the bed. Packed beds made of monosized

glass spherical particles of seven different diameters were used, as well as beds

made of 16 fractions of quartz filtration sand obtained by sieving (polydisperse

non-spherical particles). The range of bed voidages was 0.359–0.486, while the

range of bed particle Reynolds numbers was from 0.3 to 286 for spherical

particles and from 0.1 to 50 for non-spherical particles. The obtained results

were compared using a number of available literature correlations. In order to

improve the correlation results for spherical particles, a new simple equation was

proposed in the form of Ergun’s equation, with modified coefficients. The new

correlation had a mean absolute deviation between experimental and calculated

values of pressure drop of 9.04%. For non-spherical quartz filtration sand par-

ticles the best fit was obtained using Ergun’s equation, with a mean absolute

deviation of 10.36%. Surface-volume diameter (dSV) necessary for correlating the

data for filtration sand particles was calculated based on correlations for

dV = f(dm) and = f(dm).

Keywords: pressure drop, packed bed, spherical particles, quartz filtra-tion sand, non-spherical particles.

Packed beds of particles permit a widespread

means of contact between fluid and solid phases and

are used in many different industrial processes. Some

examples of their application include filtration pro-

cesses, ion-exchange, catalytic reactions, heat trans-

fer, gas scrubbing, grain drying and others. The

shape and size of particles that make up the bed are

chosen for the characteristics of the specific process.

Correspondence: T. Kaluđerović Radoičić, Faculty of Techno-

logy and Metallurgy, University of Belgrade, Karnegijeva 4, Bel-

grade, Serbia. E-mail: [email protected] Paper received: 6 May, 2015 Paper revised: 12 February, 2016 Paper accepted: 23 February, 2016

The particle size and shape always aim at high pro-

cess effectiveness, so a wide range of particles are

used. In some applications, like in down-flow granular

filters, polydisperse natural materials are used as the

particulate phase. When natural materials are used,

the shape of the particles is irregular and their size

falls into some granulometric interval. Differently

shaped particles pack with different degrees of bed

voidage, which results in different pressure drop

across the bed. The pressure drop through the

packed bed is one of the most important parameters

to be known for the adequate design of the process

as well as for the estimation of the capital and oper-

T. KALUĐEROVIĆ RADOIČIĆ et al.: FRICTION FACTOR FOR WATER FLOW… Chem. Ind. Chem. Eng. Q. 23 (1) 5766 (2017)

58

ating costs and sizing the pumps or fans required to

force the fluid through the bed.

The pressure gradient through packed beds has

been studied extensively and a large number of cor-

relations were proposed [1-17]. The most widely used

equation for pressure drop calculation was proposed

by Ergun [1]:

22

3 2 3

(1 ) (1 )150 1.75 f

p p

P- = U U

H d d (1)

The friction factor introduced by Ergun is:

3

2 1

pp

f

dPf

H U (2)

According to Eqs. (1) and (2), Ergun’s equation

for friction factor is:

p'

1501.75

Repf (3)

where:

p'Re

(1 )

p fd U (4)

For spherical particles, dp is the diameter of the

particles that constitute the packed bed, while for non-

spherical particles dp is usually taken to be the sur-

face-volume diameter dsv [18,19]. By analogy, all of

the correlations proposed for spherical particles can

be used for non-spherical particles using surface-vol-

ume diameter. Note that particle sphericity is defined

as:

SV

V

d

d (5)

where dV represents the volume diameter of the par-

ticle.

The other literature correlations for friction factor

in packed beds of spherical and non-spherical par-

ticles are shown in Table 1. Note that some authors

used different forms for friction factor and Reynolds

number with respect to Ergun’s definitions of fp and

p'Re , as shown in Table 1.

Table 1. Some important literature correlations for friction factor in packed beds of spherical particles

Reference Friction factor Eq. Re number range

Ergun [1]

p

1501.75

'Repf

(3) 1<Rep<2.4103

Macdonald et al. [2]

p

1801.80

'Repf

(6) -

Gibilaro et al. [3]

'

4.8p

18 (1 )0.33

Repf ; Note: ' 3 / (1 )p pf f

(7) -

Montillet et al. [4]

0.2

'p 3 0.5

p p

1 1000 6012

Re Re

c

p

Df = a

d

a = 0.061 ( < 0.39), a = 0.050 (( > 0.39); For (Dc/dp) > 50, term (Dc/dp)0.2 = 2.2

(8) 10 < Rep < 2.5103

3.8 ≤ Dc/dp ≤ 40-50

Kuerten, ref. in [5]

2'

3 0.5p p

25 1 ) 21 60.28

Re4 Repf =

(9) 0.1 < Rep < 4000

Hicks 6

1.2' 0.2

p3

(1 )6.8 Repf =

(10) 500 < Rep < 6104

Tallmadge 7

2 1.166' 1/6

p3 3p

150 (1 ) 4.2(1 )Re

Repf =

(11) 0.1 < Rep < 105

Lee and Ogawa 10

2' 2

3p p

1 12.5(1 ) 29.32 1.560.1 , where 0.352 0.1 0.275

2 Re Rep n

f = n

(12) 1 < Rep < 105

Cheng [9]

'pRe

pAM

f BM ,

2 11

3 1

p

c

dM

D,

2

2

1185 17

1

c

c p

DA

D d M

21/31 1

1.3 0.03 c

c p

DB

D d M

(13a)

(13)

-


59

Table 1. Continued

Reference Friction factor Eq. Re number range

Eisfeld and Schnitzlein [10]

2

'p

154

Rep

M Mf

B, M by Eq. (13a),

221.15( / ) 0.87p cB d D

(14) 0.01 < Rep <

< 1.76105

Reichelt [11]

2

'

150

Rep

p

M Mf

B, M by Eq. (13a),

22

1.5 / 0.88p cB d D (15) -

Zhavoronkov et al. [12]

2

'p

165.31.2

Rep

Af B ,

11

2 / (1 )c p

A BD d

(16) -

Raichura [13]

2

'pRe

pAM

f BM , M by Eq. (13a),

2

2

1036(1 ) 80( / )

1p cA d D

M

22.81 1.82( / )

1p cB d D

M

(17) -

Allen et al. [14]

3

2 cA A

18

1 Re Re

pA

pf

VP a bf =

H AU;

Coefficients a, b and c depend on particle type and packing structure.

Note:

ARe 4 ,

(1 ) 6

p p SVf

p p

V V dU

A A, p A

3 3, Re Re

4 2p Af f

(18) 75 < Rep < 3000

Nemec and Levec [15]

3/2 ' 4/3p

150 1.75

Repf

(19) 10 < Rep < 500

Singh et al. [16]

0.2 0.696 2.945 2

24.666Re exp[11.85(log ) ]V

S pvf

dPf

H U

Note:

pVRe ,

f VU d '

p pV3 / (1 ), Re Re / (1 )p Sf f

(20) 1257 < RepV < 2674

Ozahi et al. [17]

* 32 2

'pV

2761.76

2 1 ReO

VdPf =

L

Note:

* 21/ ,

2fP P U

'pVRe ,

(1 )

f VU d / ,p of f ' '

p pVRe Re

(21) 708 < RepV < 7773

The overview of the pressure drop correlations

for spherical particles shown in Table 1 is given in our

previous paper [20], together with the experimental

data for the friction factor for air flow through packed

beds of spherical glass particles at ambient and ele-

vated temperatures. The main conclusion of this

study was that the overall best fit of all our expe-

rimental data is given by Ergun’s [1] correlation, with

a mean absolute deviation of 10.90% [20].

Some authors proposed friction factor correl-

ations for non-spherical particles that included particle

sphericity directly in the equations (Table 1) [14-17].

Allen et al. [14] reviewed the use of different correl-

ations for pressure drop through packed beds of

spherical and non-spherical particles. They have

shown that the particle shape, arrangement, packing

method as well as surface roughness influence the

pressure drop significantly. The authors also con-

ducted pressure drop measurements in air-particles

system, using randomly packed beds of smooth and

rough glass spheres, wooden cubes, wooden cylin-

ders, acorns (ellipsoids), mixed smooth spheres of

different sizes and rounded and crushed rock with

equivalent diameters from 10.5 to 24.4 mm. The

range of p'Re numbers was 75-3000. Based on their

experimental work, they proposed the correlation for

non-spherical particles represented by Eq. (18), Table

1.

Nemec and Levec [15] investigated single-

-phase flow through packed bed reactors in the range

of p'Re numbers of 10 < p

'Re < 500 with dense and

loose packing of different uniformly sized spherical

and non-spherical particles (glass and Al2O3 spheres,

Al2O3 cylinders and rings as well as Ni-Mo trilobes

and quadralobes) in the size range of 1.26–3.49 mm.

The fluid used in their experiments was nitrogen at 10

bar. The authors concluded that Ergun’s equation

represents a good approximation of the fluid flow

through the packed bed of spherical particles in the

investigated p'Re range, while it under-predicts the

pressure drop over non-spherical particles under the

same conditions. For non-spherical particles they pro-


60

posed a new correlation which includes the particles

sphericity, given by Eq. (19), Table 1.

Singh et al. [16] investigated the pressure drop

of air flow through packed bed solar energy storage

system having large sized elements of different

shapes (spheres and cubes) with dV 125-186 mm in

the range of RepV numbers from 1257 to 2674 and

sphericities from 0.55 to 1. The correlation the authors

proposed is given by Eq. (20), Table 1.

Ozahi et al. [17] investigated the pressure drop

in air-particle systems, with particle size 6–19 mm,

and the range of sphericity of 0.55 to 1. The range of

RepV numbers in their experiments was 708–7773.

The correlation for friction factor they proposed is

given by Eq. (21), Table 1.

Several authors 9-13,21 investigated the inf-

luence of Dc/dp on pressure drop in packed beds of

particles with a general conclusion that the effect of

Dc/dp is negligible for Dc/dp >10. In addition, some

recent studies investigated the possibility to extend

the use of Ergun’s equation to polydisperse particles

systems taking into account the particle size distri-

bution [22,23].

Flow through porous media was also studied in

the field of hydrology and for different applications in

civil engineering. The experimental results obtained

were used to validate the semi-empirical relations for

non-Darcy flow [24–26].

The present study was conducted in order to

investigate the optimal choice of friction factor correl-

ation for calculating the pressure drop for water flow

through packed beds of spherical and non-spherical

particles. The experimental evaluation of literature

correlations was conducted by measuring the pres-

sure drop across packed beds of different packing

and for different water flow rates. The spherical par-

ticles used were monosized glass beads, while the

non-spherical material used was polydisperse quartz

filtration sand. The values of the equivalent diameter

and the sphericity of the quartz sand particles needed

for the calculations were obtained by using the cor-

relations for volume diameter and sphericity as a

function of mean sieve diameter proposed in our

previous paper [27]. These correlations were derived

for polydisperse fractions of quartz filtration sand with

sieve diameters in the dm interval 0.359 to 2.415 mm

[27]:

1.0787 0.0355V md d (22)

0.7942 0.063 md (23)

where dV and dm are in mm.

EXPERIMENTAL APPARATUS

The experiments were performed in the water-

–particle system schematically shown in Figure 1. The

packed bed column (f) was used for pressure drop

measurements. It was equipped with a distributor and

the calming section (e) in order to ensure the uniform

flow of water through the bed. The upwards water

flow was induced using a pump (b) and the flow rate

was measured using an electromagnetic flow meter

(d). The packed bed bulk temperature was measured

using the temperature indicator (TI). The pressure

drop in packed beds of different particles was mea-

sured using piezometers (h). The experiments were

performed with two types of particles: glass spherical

particles and polydisperse quartz filtration sand non-

-spherical particles, as shown in Figure 2. The fluid

used was deaerated water at a nearly constant tem-

perature of 20 C. In each run, water temperature was

recorded and water density and viscosity were cal-

culated. The particle characteristics and range of the

experimental conditions are summarized in Table 2.

Figure 1. Schematic diagram of the experimental system

(a-reservoir; b-pump; c-valve; d-electromagnetic flow meter;

e-calming section; f-column; g-overflow; h-piezometers;

TI-temperature indicator).

Seven kinds of mono-sized spherical glass par-

ticles were used. The experiments with glass spher-

ical particles were conducted in two cylindrical col-

umns: first column of the diameter of 40 mm was

used for 0.840–3.020 mm particles and the second

column of the diameter of 62 mm for 4.140–6.180 mm

particles. The ratio of the column diameter to the par-

ticle diameter (geometric aspect ratio) in the experi-


61

ments was between 10.0 and 47.6. As the wall effects

were not studied experimentally in this paper, it

should be noted that the literature on the range of

geometric aspect ratio in which the wall effects are

negligible is somewhat divided. Generally, the wall

effects are considered negligible at Dc/dp ratios less

than 10, but there are some researchers who found

wall effects to be significant at Dc/dp ratios as high as

15-20 [10,21]. As in the case of some of our expe-

riments the Dc/dp ratios were in that range, the exist-

ence of wall effects cannot be excluded.

The measurements of the pressure drop in the

beds of non-spherical particles were conducted using

16 fractions of quartz filtration sand obtained from the

company ”Kaolin”-Valjevo. The raw material was first

washed by fluidization to eliminate fine dust, then

dried and sieved through a number of standard sieves

with sieve openings ranging from 2.830 to 0.297 mm.

The obtained fractions had the sieve diameters in the

interval of dm = (ds,n + ds,n+1)/ 2 = 0.359 to 2.415

mm and the ratio between the two successive sieve

sizes dR =ds,n/ds,n + 1 was in the interval 1.132 to

1.715, where ds,n is the size of the opening of the

sieve through which the particle had passed

and ds,n+1 is the size of the opening of the sieve on

which the particle was retained, as shown in Table 2.

Table 2. Particle characteristics and the range of the experimental conditions

dp or dm in mm ds,n+1 / mm ds,n / mm Dc / mm p / kg m–3 U / cm s–1

Spherical particles

6.180 – – 62 2521 0.359-0.366 0.221-2.576

5.040 – – 62 2504 0.372-0.447 0.092-2.576

4.140 – – 62 2514 0.370-0.442 0.097-3.163

3.020 – – 40 2465 0.360-0.423 0.135-2.950

2.120 – – 40 2461 0.366-0.417 0.079-2.576

1.120 – – 40 2895 0.366-0.423 0.027-2.576

0.840 – – 40 2875 0.367-0428 0.018-0.448

Non-spherical particles

2.415 2.000 2.830 64 2638 0.464 0.070-1.860

1.800 1.600 2.000 64 2638 0.458 0.181-1.166

1.700 1.400 2.000 64 2638 0.453 0.098-1.036

1.583 1.166 2.000 64 2638 0.475 0.088-1.145

1.545 1.410 1.680 64 2638 0.457 0.090-1.290

1.500 1.400 1.600 64 2638 0.441 0.078-0.933

1.283 1.166 1.400 64 2638 0.449 0.067-0.943

1.201 0.991 1.410 64 2638 0.500 0.046-1.178

1.098 1.030 1.166 64 2638 0.461 0.052-0.984

1.000 0.750 1.250 64 2638 0.426 0.037-0.423

Figure 2. Some of the particles used in the experiments.


62

Table 2. Continued

dp or dm in mm ds,n+1 / mm ds,n / mm Dc / mm p / kg m–3 U / cm s–1

Non-spherical particles

0.940 0.850 1.030 64 2638 0.468 0.067-0.788

0.781 0.711 0.850 64 2638 0.470 0.015-0.626

0.656 0.600 0.711 64 2638 0.464 0.017-0.254

0.560 0.519 0.600 64 2638 0.471 0.033-0.298

0.505 0.420 0.589 64 2638 0.483 0.011-0.321

0.359 0.297 0.420 64 2638 0.486 0.011-0.091

A total of 23 runs were conducted (7 with spher-

ical particles and 16 with non-spherical particles) and

a total of 725 data points were collected (511 for

spherical and 214 for non-spherical particles). The

bed particle Reynolds number, p'Re , varied between

0.3 and 286 for spherical particles and between 0.1

and 50 for non-spherical particles. All of the water

superficial velocities used in the experiments were

below the minimum fluidization velocity for the res-

pective particles. The velocities were in the range of

0.03-0.94 UmF, where UmF represents minimum fluid-

ization velocity.


The results of the friction factor fp vs. p'Re

obtained by the experimental measurements of pres-

sure drop are shown in Figures 3 and 4 for spherical

and non-spherical particles, respectively. The com-

parison between the experimental results and the sel-

ected literature correlations is given in Table 3 and in

Figures 5 and 6. The literature correlations for spher-

ical particles shown in Table 1 were tested for all the

experimental data. The surface-volume diameter and

the sphericity needed for the calculations for non-

spherical particles were obtained using the correl-

ations from our previous paper [27] for quartz filtration

sand (Eqs. (22) and (23)). For quartz filtration sand

beds, the literature correlations specifically defined for

non-spherical particles were also tested: Allen et al.

[14], Nemec and Levec [15], Singh et al. [16] and

Ozahi et al. [17] correlations.

The mean absolute deviation between the mea-

sured values of the pressure gradient and the values

obtained from the literature correlations were cal-

culated according to the following equation:

calc measured

measured1

( / ) ( / )1

( / )

NP H P H

N P H

(24)

where N is the number of data points, (P/H)calc and

(P/H)measured are the calculated and the measured

pressure gradients.

Figure 3. fp vs. p'Re for spherical particles.

Figure 4. fp vs. p'Re for non-spherical particles.

As can be seen from Table 3, the best fit of our

experimental data for pressure drop in beds of spher-

ical particles was obtained using Cheng [9] correl-

ation, with mean absolute deviation of 10.89%. The

correlations of Macdonald et al. [2] and Montillet et al.


63

[4] also gave very good results in fitting our expe-

rimental data with mean errors of 12.18 and 13.13%,

respectively. A number of other correlations tested

gave results with mean absolute deviation in the

range of 16-20%. It should be noted that Hicks [6]

correlation gave the mean error of 50.19% for spher-

ical particles and 84.34% for non-spherical particles.

The reason for such a large error is that the range of

Rep numbers in this paper was below 183, while the

specified range of applicability of Hicks correlation is

Rep > 500.

Table 3. Comparison of experimental data for friction factor, σ

(%), with different correlations from the literature

Reference Particles

Spherical Non-spherical

Ergun [1] 19.82 10.36a

Macdonald et al. [2] 12.18 21.28 a

Gibilaro et al. [3] 18.19 13.28 a

Montillet et al. [4] 13.13 45.53 a

Kuerten, ref. in [5] 29.36 24.85 a

Hicks [6] 50.19 84.34 a

Tallmadge [7] 16.04 11.19 a

Lee and Ogawa [8] 18.61 25.30 a

Cheng [9] 10.89 31.74 a

Eisfeld and Schnitzlein [10] 18.78 11.31 a

Raichura [13] 39.07 111.39 a

Reichelt [11] 20.02 10.64 a

Zhavoronkov et al. [12] 19.04 13.58 a

Allen et al. [14] – 59.40 b

Nemec and Levec [15] – 63.52 c

Singh et al. [16] – 85.22

Ozahi et al. [17] – 78.53

This paper (Eq.(25)) 9.04 –

aUsing dSV; bUsing coefficients for crushed rock, dSVn = 24.4 mm; cUsing

coefficients for cylindrical particles

The correlations with mean absolute deviation

between experimental and correlated pressure drop

less than 20% for beds of spherical particles are

shown in Figure 5. The correlations shown in Figure 5

were calculated for Dc/dp = 25 and bed porosity of

= 0.40 (the mean values in our experiments) in

order to be able to show the correlations with direct

dependence on as lines.

Compared to the data of our previous paper

[20], in which air-spherical particles system was

investigated, the results are in the similar range for

Macdonald et al. [2] (12%) and Cheng [9] (11 and

12%) correlations, which gave very good results at

ambient temperature both for air-particles and water-

particles systems. On the other hand, the correlations

of Ergun [1], Tallmadge [7], Reichelt [11], Gibilaro et

al. [3], Einsfeld and Schnitzlein [10] and Zhavoronkov

et al. [12] performed better in air-particles system,

while the correlation of Montillet et al. [4] performed

better in water-particles systems.

Figure 5. Comparison of experimental data of fp vs. p'Re with

chosen correlations for spherical particles.

In order to improve the correlation results for

spherical particles, a new simple equation is pro-

posed in the form of Ergun’s equation, with modified

coefficient:

'

2091.75

Rep

p

f (25)

The mean absolute deviation between the

values calculated from Eq. (25) and the experimental

data is 9.04%. The new correlation is shown in com-

parison to the experimental data in Figure 3.

For non-spherical particles (quartz filtration

sand), the best fit of the experimental data was

obtained using Ergun’s equation [1], with mean abs-

olute deviation of 10.36%. The correlations of Reich-

elt [11], Tallmadge [7] and Einsfeld and Schnitzlein

[10] also gave very good results in fitting the experi-

mental data with mean errors of 10.64, 11.19 and

11.31%, respectively. The correlations with mean

absolute deviation of pressure drop less than 20% for

quartz filtration sand packed beds are shown in Fig-

ure 6. The correlations shown in Figure 6 were cal-

culated for Dc/dp = 73 and bed porosity of = 0.40

(the mean values in our experiments) in order to be

able to show the correlations with direct dependence

on as lines.


64

Figure 6. Comparison of experimental data of fp vs. p'Re with

chosen correlations for non-spherical particles.

The correlations specifically derived for non-

spherical particles gave very poor results in correl-

ating our experimental data. The mean absolute devi-

ations for these correlations were in the range of

59.40 to 85.22%. The reason for this is that the correl-

ations of Allen et al. [14], Sing et al. [16] and Ozahi et

al. [17] were derived from experiments in systems in

which the particles were of large diameters (6–20 mm

equivalent diameters) and the Rep numbers were

larger than 75, while in our system, the Rep numbers

of non-spherical particles were smaller than 25. Allen

et al. [14] and Singh et al. [16] correlations were

derived for packed bed solar energy storage systems

with particle materials being wooden cubes, crushed

rock, concrete and masonry bricks. On the other

hand, the correlation of Nemec and Levec [15] was

derived for the particles in the range of 1.26–3.49 mm,

but in the high pressure system (5–20 bar).

The very poor performance of the correlations

derived for the non-spherical particles in our quartz

sand packed beds emphasizes the fact that the

friction factor in packed beds strongly depends on

other variables besides the sphericity of the particles.

It can be concluded that the friction factor is very sys-

tem specific and that it depends on particle shape and

size, voidage, arrangement, packing method as well

as surface roughness.

It is interesting to note that almost all of the cor-

relations that gave good results in correlating our exp-

erimental data both for spherical and non-spherical

particles were in the form of Ergun’s equation, i.e.,

represented a modification of this equation with differ-

ent coefficients. From this fact it can be concluded

that the form of Ergun’s equation with two added

terms describing viscous and inertial effects is ade-

quate for representing the friction factor in packed

beds. However, the coefficients in the equation are

very system-specific and care should be taken when

choosing the adequate equation for pressure drop

calculation for the specific purpose. The correlation

chosen for the specific system should be obtained

from the data in the similar range of experimental con-

ditions as the system it is intended to be applied to.

CONCLUSIONS

The present study was conducted in order to

investigate the optimal choice of friction factor correl-

ation for water flow through packed bed of particles.

The best fit of experimental data for spherical par-

ticles was obtained using the Cheng [9] correlation

(mean absolute deviation of 10.89%). Ergun’s equa-

tion [1] gave better results in correlating the data for

non-spherical particles with mean absolute deviation

of 10.36% compared to 19.82% for spherical glass

particles. Ergun’s equation was modified in order to

improve the fit for spherical particles and a new cor-

relation was proposed. The mean absolute deviation

between the experimental data and the proposed

correlation is 9.04%.

The correlations specially derived for non-spher-

ical particles gave very poor results in correlating our

experimental data probably because they were

derived for systems with much larger particles packed

in a different arrangement. Most of the correlations

that gave good results in correlating experimental

data both for spherical and for non-spherical particles

were in the form of Ergun’s equation with modified

coefficients, thus showing that the form of Ergun’s

equation with two added terms describing viscous

and inertial effects is adequate for representing the

friction factor in packed beds. However, the coef-

ficients in the equation are very system-specific.

Acknowledgment

Financial support of the Serbian Ministry of

Education, Science and Technological Development.

(Project ON172022) is gratefully acknowledged.

Nomenclature

a empirical coefficient in Eq. (6), dimensionless

Ap area of a particle

b empirical coefficient in Eq. (6), dimensionless

c empirical coefficient in Eq. (6), dimensionless

dm particle sieve diameter

dp packed bed particle diameter


65

ds sieve opening

dSV particle surface-volume diameter

ds,n the size of the opening of the sieve through

which the particle had passed

ds,n+1 the size of the opening of the sieve on which

the particle was retained

dV particle volume diameter

Dc column diameter

fA friction factor defined according to Eq. (18)

fo friction factor defined according to Eq. (21)

fS friction factor defined according to Eq. (20)

fp friction factor defined according to Eq. (2)

f’p modified friction factor defined according to

Eq. (7)

H bed height

P pressure

P bed pressure drop, Pa

P dimensionless bed pressure drop, Eq.(21)

Rep =(fUdp)/μ particle Reynolds number

p'Re =(fUdp)/(μ(1-ε)) bed particle Reynolds num-

ber

RepV particle Reynolds number defined according

to Eq. (20)

RepV’ bed particle Reynolds number defined

according to Eq. (21)

ReA bed particle Reynolds number defined

according to Eq. (18)

U superficial fluid velocity

UmF minimum fluidization velocity (superficial)

Vp volume of a particle

Greek letters

voidage

μ fluid viscosity

ρf fluid density

ρp particle density

σ mean absolute deviation

particle sphericity, dimensionless

Subscripts

f fluid

mF minimum fluidization

p particle

REFERENCES

[1] S.S. Ergun, Chem. Eng. Prog. 48 (1952) 89-94

[2] F. Macdonald, M.S. El-Sayed , K. Mow, F.A.L. Dullien,

Ind. Eng. Chem. Fundam. 18 (1979) 199-208

[3] L.G. Gibilaro, R. Di Felice, S.P. Waldram, Chem. Eng.

Sci. 40 (1985) 1817-1823

[4] A. Montillet, E. Akkari, J. Comiti, Chem. Eng. Process. 46

(2007) 329-333

[5] H. Watanabe, Inte. J. Eng. Fluid Mech. 2 (1989) 93-108

[6] R.E. Hicks, Ind. Eng. Chem. Fundam. 9 (1970) 500-502

[7] J.A. Tallmadge, AIChE J. 16 (1970) 1092-1093

[8] J. Lee, K. Ogawa, J. Chem. Eng. Jpn. 27 (1994) 691-693

[9] N.S. Cheng, Powder Technol. 210 (2011) 261-266

[10] B. Eisfeld, K. Schnitzlein, Chem. Eng. Sci. 56 (2001)

4321-4329

[11] W. Reichelt, Chem. Ing. Tech. 44 (1972) 1068-1071

[12] N.M. Zhavoronkov, M.E. Aerov, N.N. Umnik, J. Phys.

Chem. 23 (1949) 342-361 (in Russian)

[13] R.C. Raichura, Exp. Heat Transfer 12 (1999) 309-327

[14] K.G. Allen, T.W. von Backström, D.G. Kröger, Powder

Technol. 246 (2013) 590-600

[15] D. Nemec, J. Levec, Chem. Eng. Sci. 60 (2005) 6947-6957

[16] R.Singh, R.P. Saini, J.S. Saini, Sol. Energy 80 (2006)

760-771

[17] E.Ozahi, M.Y.Gundogdu, M.Ö. Carpinlioglu, Adv. Powder

Technol. 19 (2008) 369-381

[18] D. Geldart, Powder Technol. 60 (1990) 1-13

[19] W.-C. Yang, in: Handbook of Fluidization and Fluid-par-

ticle Systems, W.-C. Yang, Marcel Dekker, New York,

2003

[20] R. Pešić, T. Kaluđerović Radoičić, N. Bošković-Vra-

golović, Z. Arsenijević, Ž. Grbavčić, Chem. Ind. Chem.

Eng. Q. 21 (2015) 419–427

[21] R. Di Felice, L.G. Gibilaro, Chem. Eng. Sci. 59 (2004)

3037-3040

[22] M. Mayerhofer, J. Govaerts, N. Parmentier, H. Jeanmart,

L. Helsen, Powder Technol. 205 (2011) 30-35

[23] A. Luckos, J.R. Bunt, Fuel 90 (2011) 917-921

[24] M. Sedghi-Asl, R. Hassan, J. Hydrol. Eng. 49 (2011) 814-

–817

[25] K.N. Moutsopoulos, I.N. Papaspyros, V.A. Tsihrintzis, J.

Hydrol. 374 (2009) 242-254

[26] M.B. Salahi, M. Sedghi-Asl, M. Parvizi, J. Hydrol. Eng.

(2015) 04015003

[27] T. Kaluđerović Radoičić, M.Đuriš, R.Garić-Grulović, Z. Ar-

senijević, Ž.Grbavčić, Powder Technol. 254 (2014) 63-71.


66

TATJANA KALUĐEROVIĆ RADOIČIĆ1

NEVENKA

BOŠKOVIĆ-VRAGOLOVIĆ1

RADMILA GARIĆ-GRULOVIĆ2

MIHAL ĐURIŠ2

ŽELJKO GRBAVČIĆ1

1Tehnološko-metalurški fakultet,

Univerzitet u Beogradu, Karnegijeva 4,

Beograd, Srbija 1Institut za hemiju, tehnologiju i

metalurgiju, Univerzitet u Beogradu,

Njegoševa 12, Beograd, Srbija

NAUČNI RAD

KOEFICIJENT TRENJA FLUID-ČESTICE U PAKOVANIM SLOJEVIMA SFERIČNIH I NESFERIČNIH ČESTICA

Cilj ovog rada je bio eksperimentalno ispitivanje koeficijenta trenja fluid-čestice prilikom

strujanja vode kroz pakovani sloj sferičnih i nesferičnih čestica na sobnoj temperaturi. U

eksperimentima je meren pad pritiska prilikom strujanja fluida kroz pakovani sloj. Na osno-

vu dobijenih rezultata, izvršena je evaluacia različitih literaturnih korelacija. U eksperimen-

tima je korišćeno sedam vrsta monodisperznih sferičnih staklenih čestica, kao i 16 frakcija

polidisperznih nesferičnih čestica filtracionog peska različitih dimenzija, dobijenih proseja-

vanjem. Opseg poroznosti pakovanih slojeva je bio od 0,359 do 0,486, dok je opseg vred-

nosti Rejnlodsovog broja za čestice bio od 0,3 do 286 za sferične čestice i od 0,1 do 50 za

nesferične čestice. Dobijeni rezultati su korelisani korišćenjem većeg broja literaturnih

korelacija. U cilju poboljšanja rezultata korelisanja za sferične čestice, predložena je nova

jednačina u formi Ergunove jednačine sa modifikovanim koeficijentima. Srednje apsolutno

odstupanje eksperimentalnih od izračunatih vrednosti za predloženu korelaciju iznosilo je

9,04%. Za nesferične čestice kvarcnog filtracionog peska, najbolji rezultati su dobijenu

korišćenjem Ergunove jednačine, sa srednjim apsolutnim odstupanjem od 10,36%. Povr-

šinsko-zapreminski prečnik (dSV) koji je neophodan za korelisanje eksperimentalnih poda-

taka za nesferične čestice je računat na osnovu korelacija za dV = f(dm) i = f(dm) koje su

predložene u našem prethodnom radu [27].

Ključne reči: pad pritiska, pakovani sloj, sferične čestice, kvarcni filtracioni pesak,

nesferične čestice.






67

HAI-PENG GOU1,2

GUO-HUA ZHANG1

KUO-CHIH CHOU1,2

1State Key Laboratory of Advanced

Metallurgy, University of Science

and Technology Beijing, Beijing,

China 2Collaborative Innovation Center of

Universal Iron and Steel

Technology, University of Science

and Technology Beijing, Beijing,

China

SCIENTIFIC PAPER

UDC 621.762:669.295:546.72:66

https://doi.org/10.2298/CICEQ150521007G

PREPARATION OF TITANIUM CARBIDE POWDER FROM ILMENITE CONCENTRATE

Article Highlights

• Titanium carbide powder was prepared by the carbothermic reduction of ilmenite con-

centrate

• The iron and titanium carbide in the reduction products were economically separated by

using ferric chloride solution

• The iron in the waste liquid was recycled in the form of -FeOOH

Abstract

A new process of producing titanium carbide powder from ilmenite concentrate

was put forward. The ilmenite concentrate was reduced by graphite powder at

1500 C for 6 h, with the reduction products of TiC, Fe and MgO. The porous

reduction products were agitation leached in ferric chloride solution at 25 C for

60 min. After the filtration, TiC powder accompanied by little MgAl2O4 and

Mg2SiO4 were obtained. Finally, the element Fe was recycled from the filtrate in

the form of -FeOOH after blowing air for 3 hours at 80 C. The size of rod-like

-FeOOH particle was less than 1 µm.

Keywords: titanium carbide, ferric chloride solution, ilmenite concen-trate, -FeOOH.

Titanium carbide, TiCx, existing as a homogen-

eous phase within the limits 0.47 < x < 1.0 [1], has a

NaCl-type of structure. Because of its high melting

point (3067 C), high hardness (32.4 GPa), good

chemical inertness and good electrical conductivity

(310-7 S/cm), titanium carbide has been found tre-

mendous applications in various fields, such as wear-

resistant material, cutting tool and anodes in lithium-

-ion batteries [2-4].

Nowadays, various methods have been adopted

to prepare titanium carbide, e.g., self-propagating

high temperature synthesis (SHS) [1], mechanically

activated sintering [5], thermal plasma [6], carbo-

thermic reduction [7,8], etc. Considering the abundant

reserves of ilmenite concentrate in the world, the pre-

paration of titanium carbide by ilmenite concentrate is

receiving more and more attentions. Most of the

studies were focused on the carbothermic reduction

process [9-11], while how to separate the reduction

Correspondence: G.-H. Zhang, State Key Laboratory of Adv-

anced Metallurgy, University of Science and Technology Bei-

jing, Beijing, 100083, China. E-mail: [email protected] Paper received: 21 May, 2015 Paper revised: 19 February, 2016 Paper accepted: 26 February, 2016

products of iron and titanium carbide economically is

also quite significant. After the carbothermic reduction

of ilmenite concentrate, Welham and Williams [7]

used 3% HCl to leach the products for 24 h at room

temperature. As a result, almost all of the iron went

into the waste liquid and could not be recycled. On

the other hand, there are many reports on the pro-

duction of TiC reinforced iron-based composite from

ilmenite [12-16]. After being milled in a planetary ball

mill, TiC reinforced iron-based composite was syn-

thesized via microwave heating [12,13], electric dis-

charge assisted mechanical milling (EDAM) [14] or

carbothermic reduction [15,16]. It is economical to

convert raw materials directly to composite materials.

In this paper, a new process of separating iron

and titanium carbide via lixiviation was put forward.

The ferric chloride solution was used for leaching to

separate reduction productions of iron and titanium

carbide. After leaching, the main component of the fil-

trate was ferrous chloride. The iron in the filtrate could

be recycled in the form of -FeOOH [17,18]. -FeOOH

could be used as an important by-product to produce

ultrafine Fe2O3 [19]. After recycling the elemental Fe,

the main component of the rest filtrate was ferric

H.-P. GOU et al.: PREPARATION OF TITANIUM CARBIDE POWDER… Chem. Ind. Chem. Eng. Q. 23 (1) 6772 (2017)

68

chloride, which could be reused for the next leaching

process.

EXPERIMENTAL

The ilmenite concentrate, produced in Panz-

hihua, Sichuan, China, was examined by X-ray fluor-

escence (XRF) and X-ray diffraction (XRD). The

results of XRF and XRD are presented in Table 1 and

Figure 1a, respectively. The main mineral phase of the

ilmenite concentrate is FeTiO3, while the main impur-

ity element is Mg, which exists in solid-solution of

(Mg,Fe)(Ti,Fe)O3 [20]. As shown in Figure 1b, it should

be noted that the peaks of FeTiO3 and (Mg,Fe)(Ti,Fe)O3

are similar and overlapped in the XRD pattern

because of the same crystalline structure.

Figure 1. XRD patterns of the ilmenite concentrate.

According to the authors’ previous study [10],

the optimized carbothermic reduction parameters

were established as: molar ratio of C to FeTiO3 4:1,

reduction temperature 1500 C and reduction time 6

h. The main reactions occurred during the carbo-

thermic reduction are shown in Eqs. (1)-(4). All the

standard Gibbs energy changes, G , are calculated

by Factsage 6.4:

3FeTiO 4C Fe TiC 3CO

G = –497.7T + 711201.1 (1)

3MgTiO 3C MgO TiC 2CO

G = –342.2T + 558842.7 (2)

2 3 2 4MgO Al O MgAl O

G = –6.9T + 22840.7 (3)

2 2 42MgO SiO Mg SiO

G = –4.4T + 65772.6 (4)

The ilmenite concentrate and graphite powder

(Sinopharm Chemical Reagent Co., Ltd, Chemical

Pure, 98%) were mixed uniformly in an agate mortar

(Changzhou Putian Instrument Manufacture CO., Ltd,

diameter 105 mm). Then the mixtures were made into

cylindrical briquettes with the addition of PVA (2 wt.%).

The diameter and weight of the cylindrical briquettes

were 18 mm and 2 g, respectively. When the tempe-

rature of the vertical tube furnace reached 1500 C,

the alumina crucible with the briquettes was put into

the furnace under a protective argon gas atmosphere

(0.8 L/min). After reacting for 6 h, the crucible was

taken out of the furnace quickly and cooled by the

argon stream (1.5 L/min). The reduction products

were examined by XRD and scanning electron micro-

scope (SEM) to investigate their phase compositions

and microstructure.

The reduction products were agitation leached

by ferric chloride solution (0.5 mol/L) at 25 C in an

electro-thermostatic water bath. The stirring rate was

200 rpm and the pulp density for each experiment

was 20 g/L. The leaching time was 1, 2, 5, 10, 20, 30

and 60 min, respectively. After leaching, the solid

phase was separated from the liquid phase by suction

filtration as soon as possible. After being rinsed

thoroughly by deionized water, the obtained solid

phase was examined by XRD and SEM.

Table 1. Chemical compositions of ilmenite concentrate (wt.%)

Component FeO TiO2 SiO2 CaO Al2O3 MgO SO3

Content 39.30 43.68 3.15 1.28 2.91 7.99 0.62

Component Na2O MnO Cr2O3 ZnO P2O5 In2O3 Total

Content 0.28 0.69 0.03 0.02 0.03 0.02 100


69

The filtrate after leaching was collected in a

beaker and then heated with blowing air (0.5 L/min).

The temperature was set to be 80 C and was not

changed during this process. The purpose of blowing

air into the filtrate was to accelerate the reaction. After

3 h, a yellow solid precipitate appeared. The preci-

pitate was collected by suction filtration and examined

by XRD and field emission scanning electron micro-

scope (FE-SEM).


Carbothermic reduction of the ilmenite concentrate

After 6 h of the carbothermic reduction, the per-

centage of total mass loss ratio was 41.88, which was

consistent with the maximum theoretical mass loss

ratio [10]. It illustrates that the carbothermic reduction

reacted completely. The XRD patterns of the red-

uction products are presented in Figure 2, from which

it can be concluded that the reduction products were

mainly made up of TiC, Fe and MgO. The morphology

images of the ilmenite concentrate and the reduction

products morphology are shown in Figure 3a and b,

respectively. Compared with the ilmenite concentrate,

the reduction products were porous, which was due to

the generation and evolution of CO gas. Backscat-

tered electron (BSE) image of the main phases inside

the reduction products is displayed in Figure 3c,

which indicates that there are three different regions

in the samples. Based on the results, the energy dis-

persive spectrometer (EDS) analyses performed at

different regions are shown in Table 2. It is obvious

that the main impurity element Mg mostly existed in

the form of MgO after the carbothermic reduction, which

is consistent with the results of the XRD analyses.

Purification of titanium carbide powders

The percentages of mass loss with different

leaching time are shown in Figure 4a. After leaching

for 10 min, the mass loss vs. time curve reached a

plateau of 47.6%. As shown in Figure 3b, the porous

structure of the reduction products was advantageous

to the leaching process. XRD pattern of the leaching

products after leaching for 60 minutes is presented in

Figure 4b. Compared with Figure 2, the peaks of Fe

and MgO disappeared in Figure 4b. Based on Figures

2 and 4b, the following reactions occurred in the pro-

cess of lixiviation. The hydrogen ions in Eq. (6) were

resulted from the hydrolysis of ferric chloride solution:

3 22FeCl Fe 3FeCl (5)

222H MgO Mg H O (6)

Figure 2. XRD patterns of the reduction products.

Table 2. EDS results of different phases in Figure 3c

Phase Elements mass fraction

Fe 88.75%Fe; 5.99%C; 2.37%Si; 2.89%Ti

TiC 85.43%Ti; 13.85%C; 0.71%V

MgO 30.80%O; 53.68%Mg; 1.91%Ca; 9.80%Si; 0.78%Fe;

3.02%Ti

Figure 3. a) SEM morphology image of the ilmenite concentrate; b) SEM image of the reduction products; c) BSE image of the mainly

compositions inside the reduction products.


70

After leaching, the main phases were TiC,

MgAl2O4 and Mg2SiO4. BSE images of the leaching

products are shown in Figure 5. The overall distri-

bution of the leaching products is displayed in Figure

5a. Most of the iron was removed by ferric chloride

solution. As a result, TiC particles which were integ-

rated with Fe tightly were scattered into individual

particles. The wetting angle between TiC and Fe at

1500 C is 30 [21]. As shown in Figure 5b, there was

still a little bit of Fe inside TiC particles. This part of

the iron was wrapped by TiC particles and could not

be removed by ferric chloride solution. From Figure

5c and d, it could be seen that a few impurities of

MgAl2O4 and Mg2SiO4 were combined with TiC par-

ticles. EDS analyses performed at different phases

are shown in Table 3, which were in agreement with

the XRD patterns in Figure 4b. The removal of MgAl2O4

and Mg2SiO4 are complicated and are now in prog-

ress. Or, the mixture of TiC, MgAl2O4 and Mg2SiO4

may be considered as a new TiC based composite

material since all of them have very high hardness.

Recycle of the waste liquid

When blowing air into the filtrate, the ferrous

chloride solution was oxidized to ferric chloride sol-

ution, and a yellow solid precipitate appeared. The

Figure 4. a) The percentages of mass loss with different leaching time; b) XRD patterns of the leaching products

with leaching for 60 min.

Figure 5. BSE images of the leaching products.


71

XRD patterns of the precipitate are presented in

Figure 6a, which indicates that the precipitate is aka-

ganeite (-FeOOH). As seen from the FE-SEM mor-

phology image of -FeOOH shown in Figure 6b,

-FeOOH with the length less than 1 m was rod-like

shaped, which was consistent with the reported expe-

riment phenomena [22,23]. The rod-like particles of

-FeOOH could be dissolved and reprecipitate as

hematite (-Fe2O3). Under different conditions, -Fe2O3

could be in different morphologies of spheres, cubes

or double ellipsoids [22]. Therefore, -FeOOH is a

valuable by-product. The main reaction occurring dur-

ing blowing air into the waste liquid is shown in Eq.

(7):

2 2 2 312FeCl 3O 2H O 8FeCl 4 -FeOOH (7)

Table 3. EDS results of different phases in Figure 5

Phase Elements mass fraction

Fe 87.94%Fe; 9.94%C; 2.12%Ti

TiC 86.73%Ti; 12.64%C; 0.63%V

MgAl2O4 48.84%O; 15.84%Mg; 31.67%Al; 3.65%Ti

Mg2SiO4 45.40%O; 34.69%Mg; 17.34%Si; 1.02%Ca; 1.55%Ti

As can be concluded from Eq. (7), 1/3 Fe con-

tent is recovered in the form of FeOOH and the

remaining Fe still existed as ferric chloride in solution,

which could be reused in the next leaching process.

However, as the accumulation of the elemental Mg in

the new ferric chloride solution, the recycle of the

ferric chloride solution would be terminated when

MgCl2 reached saturation.

CONCLUSION

The purification of titanium carbide powder from

ilmenite concentrate was investigated in this article.

After the carbothermic reduction at 1500 C for 6 h

and lixiviation by ferric chloride solution at 25 C for

10 min, TiC powder with little MgAl2O4 and Mg2SiO4

was produced. The mass loss ratios during the carbo-

thermic reduction and lixiviation process were 41.88

and 47.6%, respectively. The carbothermic reduction

product Fe went into the waste liquid and was

recycled in the form of -FeOOH after blowing air for

3 h at 80 C. The waste liquid was oxidized to ferric

chloride solution for the next process of lixiviation.

Acknowledgements

Thanks are given to the financial supports form

the Fundamental Research Funds for the Central

Universities (FRF-TP-15-009A3) and the National

Natural Science Foundation of China (51474141).

REFERENCES

[1] J.B. Holt, Z.A. Munir, Combustion synthesis of titanium

carbide: theory and experiment. J. Mater. Sci. 21 (1986)

251-259

[2] G.E. Hollox, R.E. Smallman, J. Appl. Phys. 37 (1966)

818-823

[3] N. durlu, J. Eur. Ceram. Soc. 19 (1999) 2415-2419

[4] P. Patel, I.S. Kim, P.N. Kumta, Mater .Sci. Eng. B 116

(2005) 347-352

[5] M. Razavi, M. R. Rahimipour, R. Kaboli, J. Alloys Compd.

460 (2008) 694-698

[6] L. Tong, R.G. Reddy, Scripta Mater. 52 (2005) 1253-1258

[7] N.J. Welham, J.S. Williams, Metall. Mater. Trans. B 30

(1999) 1075-1081

[8] N.J. Welham, Miner. Eng. 9 (1996) 1189-1200

[9] K.S. Coley, B.S. Terry, P. Grieveson, Metall. Mater.

Trans. B 26 (1995) 485-494

[10] H.P. Gou, G.H. Zhang, K.C. Chou, Metall. Mater. Trans.

B 46 (2014) 48-56

[11] S.K. Gupta, V. Rajakumar, P. Grieveson, Metall. Trans. B

20 (1989) 735-745

Figure 6. a) XRD patterns of the precipitate; b) FE-SEM morphology image of -FeOOH.


72

[12] M. Razavi, M.S. Yaghmaee, M.R. Rahimipour, S.S.

Razavi-Tousi, Int. J. Miner. Process. 94 (2010) 97-100

[13] M. Razavi, M.R. Rahimipour, T. Ebadzadeh, S.S. Razavi-

-Tousi, Bull. Mater. Sci. 32 (2009) 155-160

[14] A. Calka, D. Oleszak, N. Stanford, J. Alloys Compd. 459

(2008) 498-500

[15] Y. Chen, Scripta Mater. 36 (1997) 989-993

[16] M. Razavi, M.R. Rahimipour, Ceram. Int. 35 (2009) 3529-

-3532

[17] C. Rémazeilles, Ph. Refait, Corros. Sci. 49 (2007) 844-

-857

[18] P. Refait, J.M.R. Génin, Corros. Sci. 39 (1997) 539-553

[19] E.A. Deliyanni, D.N. Bakoyannakis, A.I. Zouboulis, K.A.

Matis, L.Nalbandian, Microporous Mesoporous Mater. 42

(2001) 49-57

[20] R.G. Berman, L.Y. Aranovich, Contrib. Mineral. Petrol.

126 (1996) 1-24

[21] R. Warren, J. Mater. Sci. 15 (1980) 2489-2496

[22] J.K. Bailey, C.J. Brinker, M.L. Mecartney, J.Colloid

Interface Sci. 157 (1993) 1-13

[23] S. Musić, S. Krehula, S. Popović, Mater. Lett. 58 (2004)

444-448.

HAI-PENG GOU1,2

GUO-HUA ZHANG1

KUO-CHIH CHOU1,2

1State Key Laboratory of Advanced

Metallurgy, University of Science and

Technology Beijing, Beijing, China 2Collaborative Innovation Center of

Universal Iron and Steel Technology,

University of Science and Technology

Beijing, Beijing, China

NAUČNI RAD

PRIPREMA PRAŠKASTOG TITAN-KARBIDA OD KONCENTRATA ILMENITA

Razvijen je novi proces za dobijanje praškastog titanijum karbida od koncentrata ilmenita.

Koncentrat ilmenita je usitnjen grafitnim prahom na 1500 C u toku 6 h, pri čemu su pro-

izvodi redukcije TIC, Fe i MgO. Porozni redukcioni proizvodi su luženi u rastvoru feri-hlo-

rida na 25 C uz mešanje 60 min. Nakon filtracije, dobijeni su prah TiC uz malo MgAl2O4 i

Mg2SiO4. Konačno, elementarno gvožđe je reciklisano iz filtrata u obliku -FeOOH nakon

uduvavanja vazduh tokom 3 h na 80 C. Veličina štapišastih čestica -FeOOH je manja od

1 m.

Ključne reči: Titan-karbid, gvožđe-hlorid, koncentrat ilmenite, -FeOOH.






73

MILANA M. ZARIĆ1

MIRKO STIJEPOVIC1

PATRICK LINKE2

JASNA STAJIĆ-TROŠIĆ1

BRANKO BUGARSKI3

MIRJANA KIJEVČANIN3

1Institute of Chemistry, Technology

and Metallurgy, University of

Belgrade, Belgrade, Serbia 2Department of Chemical

Engineering, Texas A&M

University at Qatar, Education City,

Doha, Qatar 3Faculty of Technology and

Metallurgy, University of Belgrade,

Belgrade, Serbia

SCIENTIFIC PAPER

UDC 620.92:66.021.4


TARGETING HEAT RECOVERY AND REUSE IN INDUSTRIAL ZONE

Article Highlights

• Heat recovery and reuse of waste heat via indirect heat integration

• Increasing of energy efficiency and reducing consumption of fossil fuel

• Linear programming (LP) used for method formulation

• Industrial zone energy integration strategy

Abstract

In order to reduce the usage of fossil fuels in industrial sectors by meeting the

requirements of production processes, new heat integration and heat recovery

approaches are developed. The goal of this study is to develop an approach to

increase energy efficiency of an industrial zone by recovering and reusing waste

heat via indirect heat integration. Industrial zones usually consist of multiple inde-

pendent plants, where each plant is supplied by an independent utility system, as

a decentralized system. In this study, a new approach is developed to target

minimum energy requirements where an industrial zone would be supplied by a

centralized utility system instead of decentralized utility system. The approach

assumes that all process plants in an industrial zone are linked through the

central utility system. This method is formulated as a linear programming prob-

lem (LP). Moreover, the proposed method may be used for decision making rel-

ated to energy integration strategy of an industrial zone. In addition, the pro-

posed method was applied on a case study. The results revealed that saving of

fossil fuel could be achieved.

Keywords: heat recovery, energy efficiency, heat integration, LP formul-ation.

Global industrial growth has triggered energy

consumption levels by the industrial sectors [1].

Energy intensive processes rely on usage of fossil

fuels to provide their energy requirements. Increased

fossil fuel consumption leads to undesirable increase

in greenhouse gas (GHG) emissions, causing climate

changes [2]. There are two basic concepts to dec-

rease the use of fossil fuels: replacing the fossil fuels

with renewable energy or increasing the energy effi-

ciency of the process [3]. In order to meet the ind-

ustrial requirements while reducing emissions of GHG,

the efficiency of production processes needs to be

improved. Also, new technologies are developed to

Correspondence: M. Kijevčanin, Faculty of Technology and

Metallurgy, University of Belgrade, Karnegijeva 4, 11120 Bel-

grade, Serbia. E-mail: [email protected] Paper received: 22 June, 2015 Paper revised: 2 February, 2016 Paper accepted: 4 March, 2016

target process integration and process intensification

[4]. Moreover, process inefficiencies cause substan-

tial heat and energy loss to the environment. Bendig

et al. classified energy loss as avoidable and unavoid-

able heat loss, and waste heat was defined as avoid-

able heat loss [5]. Heat recovery of waste heat is

considered as a very promising strategy for enhanc-

ing the overall energy efficiency in a process [6].

Recovery and reuse of waste heat can be applied at

the process level as well as at the plant level [7].

Recent studies applied waste heat recovery and

reuse strategy on various levels - industries, industrial

zones and continuous casting process [8–10]. In com-

plex systems, which consist of multiple processing

plants, each plant has its own independent operating

and maintenance schedules, which sets difficulties

and limitations during integration process. Also, an

important factor to consider is the distance between a

M.M. ZARIĆ et al.: TARGETING HEAT RECOVERY… Chem. Ind. Chem. Eng. Q. 23 (1) 7382 (2017)

74

heat sink and a heat source, including heat loss due

to transport.

Process level heat recovery is usually being

referred to as a direct heat recovery, where heat rec-

overy is performed between hot and cold process

streams. This is not always achievable, due to many

practical implementation issues such as: control-

lability, possibility of generating hazards, different

operating scenarios, long distance between process

streams, etc. Plant level heat recovery helps over-

come the shortcomings of process level heat recovery

and is often defined as indirect heat recovery. This

type of integration is performed via intermediate fluids

(such as steam, hot oil, flue gas, etc.), and provides

advantages such as operational flexibility, control-

lability, as well as avoidance of hazards. However,

comparing the two levels of heat recovery, during

indirect heat recovery temperature driving forces are

reduced, which often results in lower heat recovery

and less energy saved [11]. Wang et al. considered

both direct and indirect heat integration, as well as

combination of both direct and indirect heat recovery

involving the features of both giving more design

options [12]. Recently, Miah et al. have maximized

the heat recovery of diverse production lines by com-

bining the direct and indirect heat exchange from

zonal to factory level [13].

The Pinch analysis (PA) method introduced by

Linnhoff et al. is one of the most commonly used

methods that can estimate possible heat recovery

within an individual process [14]. PA methods are

based on thermodynamic principles to determine

maximum heat recovery potential, and hence can be

used to construct efficient designs for heat exchanger

network (HEX) [15-20]. The main disadvantage of this

method is that it does not allow options for forbidden

or preferred matches between process streams.

Moreover, the total site analysis (TSA) method has

been introduced to improve heat recovery within a

given plant [21]. TSA applies energy integration

between multiple processes to enable maximum indi-

rect heat recovery potential. Processes within a plant

are considered to be supplied by a common utility

system, which provides the required heat and power.

Dhole and Linnhoff introduced TSA to establish tar-

gets for heat recovery by integrating processes and

optimizing the quantity of utility used in plants [21].

Practical implementation of TSA as an energy conser-

vation concept has been improved in previous studies

[22-24]. Chew et al. pointed out that TSA methods

should be extended to design, operability, reliability/

/availability, maintenance, regulatory policy, as well

as economics of utility systems [25]. Further studies

investigated total site heat integration considering

pressure drops and utilizing the TS Heat Integration

profiles for assessing the process modifications to

decrease capital costs [26,27]. Improvement of TSA

method has been investigated by Varbanov et al.,

who implemented a modification to enable the use of

different minimum temperature differences [28]. New

graphical approaches are proposed to present better

clarity for the quantitative interaction between the pro-

cess and utility system targets [29]. TSA methods

mainly rely on graphical techniques that cannot pro-

vide precise estimations [3,28,30]. Moreover, steam

superheating considerations are overlooked, often

causing heat recovery potential to be overestimated

[31]. Mathematical programming techniques have

been developed to overcome the aforementioned

drawbacks TSA in identifying optimal HEXs designs.

Papoulias and Grossman proposed a mathematical

programing technique that identifies an option for for-

bidden matches, whereas Becker and Marachel deve-

loped a mathematical programming model by adding

an option for intermediate heat transfer units [32,33].

Liew et al. developed an extended methodology TS

Heat Integration in a steam system that considers the

water sensible heat (boiler feed water preheating and

steam superheating during steam generation) using a

systematic numerical tool [34]. Several reviews of rel-

evant publications on heat exchange synthesis and

process integration have been published [35,36].

Recent studies have introduced an optimal design

approach and multi-objective optimization methods of

cogeneration systems based on exergo-economic

and exergo-environmental parameters [37,38].

The goal of this study is to develop an approach

to maximize energy efficiency of an industrial zone

through waste heat recovery and reuse, via indirect

heat integration. Previous work done by Stijepovic

and Linke proposed a method targeting maximal

waste heat recovery and reuse across decentralized

utility system [9]. This method is based on a study of

an industrial zone consisting of independent plants

operating multiple processes. It is considered that

each plant is served by an independent utility system

and each plant has been optimized for energy effi-

ciency. In this study, a new approach is developed to

target minimum energy requirements in an industrial

zone supplied by a centralized utility system instead

of decentralized utility system. To reveal potential

waste heat streams, the concept of exergy was used,

a method described by Stijepovic and Linke [9]. The

transshipment model is adopted to estimate maximal

heat recovery from recognized waste heat streams.

The study is based on targeting maximum waste heat


75

recovery potentials in a centralized utility system prior

to the design of optimal network. Furthermore, the

proposed approach may assist in the decision making

process regarding the retrofitting strategy for utility

system configuration in an industrial zone. For

example, it can reveal whether introducing a central-

ized utility system between multiple plants is justified.

Problem definition

A plant usually comprises several processes,

where a common utility system provides overall heat

for all processes. In this study, an industrial zone

consists of multiple independent plants and it is con-

sidered that all plants are served by one centralized

utility system, which provides required heat and

energy for all plants, instead of each plant having a

separate utility system. Both initial and target tempe-

ratures, heat capacities and heat loads are specified

for each hot and cold process stream. Centralized

utility system uses fossil fuel to generate very high

pressure steam (VHP), where thermodynamic state is

reduced to required utilities pressures and tempera-

tures by let-down stations. Depending on the require-

ments of the process, different types of high pressure

(HP), medium pressure (MP), and low pressure (LP)

utilities are generated. Heat demands for required

utilities define the overall consumption of fossil fuel.

Generally speaking, two types of process

streams exist: 1) hot process streams have to be

cooled down to a specified temperature, and 2) cold

process streams have to be heated to a specified

temperature. Hot process streams are cooled down

using cold process streams. Similarly, cold process

streams are heated up using hot process streams.

Any hot or cold process steams which are unable to

reach their specified temperature solely via heat

exchange must be cooled down or heated up addi-

tionally by using external cold or hot utilities, respect-

ively. Any excess heat that is released to cold utilities

may then be used to provide heat for another plant

within the industrial zone. Therefore, excess heat can

be used to generate utilities, which can then be used

as a heat source in another plant within the industrial

zone [31].

Process streams that eject excess heat into cold

utilities can be identified as a potential heat source,

as shown in Figure 1. Cold utilities mainly use air or

cooling water to cool down a hot process stream.

Heat released to cold utilities is considered waste

heat (Figure 1a), which can later be used as a heat

source to generate utilities subsequently referred to

as “recoverable utility” (Figure 1b). The role of intro-

duced recoverable utilities in heat integration is to

transfer heat from a process where excess heat is

identified to a process with heat deficit. The generat-

ed recoverable utility replaces required hot utilities,

either totally or partially, leading to decrease in overall

industrial zone heat demands set prior to heat integ-

ration.

Figure 2 illustrates a utility system with indirect

heat integration between processes within an indus-

trial zone. The utility system generates VHP steam

that is converted by let-down stations to hot utilities

HP steam, MP steam, and LP steam. Recoverable

utilities are generated using excess heat from hot

process streams and are directed to hot utility steam

headers, where they are linked with the specified

steam from let-down station. This leads to decrease

in demands of generating VHP as well as the usage

of fossil fuels.

Model formulation

The proposed model is depicted in the heat cas-

cade diagram in Figure 3. Hot streams represent heat

sources, and recoverable utilities represent heat sinks

during heat exchange which is carried out in tempe-

rature intervals k. Temperature intervals account for

the thermodynamic constrains that control heat trans-

fer in order to guarantee feasible heat transfer con-

ditions in each interval. This has been ensured by

partitioning the entire temperature range into small

temperature intervals, which are defined by initial and

final temperature of present streams. The entire

range of temperature values is set in decreasing

order in the cascade diagram. There are k tempera-

ture intervals for k+1 values of temperature, each

represented by a separate block of heat exchange

between sources (hot process streams) and sinks

Figure 1. a) Hot process stream cooled by cold utility; b) hot process stream generating recoverable utility.


76

(recoverable utilities). Index sets, parameters and

variables that are required by the problem are defined

in the nomenclature.

Each hot process stream with an excess heat

content at temperature interval, k, Qi,k, can exchange

heat content with any recoverable utility, j, that is pre-

sent in that particular temperature interval, k. The

transferred heat from hot process stream i to recover-

able utility j in temperature interval k is represented by

Qi,,j,k. Part of the heat content of hot process stream

that has not been exchanged in the temperature inter-

val k, is transferred to the next, lower temperature

interval, k+1, as a heat residual, Ri,k (Figure 3).

The proposed approach is defined as follows:

VHPminOF m (1)

Figure 2. Utility system with indirect heat integration.

Figure 3. Heat cascade diagram.


77

Set to constrains:

, , 1 , , ,

1

JH

i k i k i j k i k

j

R R Q Q , ,i Hk k TI (2)

, , ,

1

IC

i j k j k

i

Q Q , ,j Ck k TI (3)

req 0jm m , req ,RHU j Ck (4)

1 1

REQ JVHP

req j

req j

m m m (5)

, , ,, , 0i k i j k jR Q m (6)

,0 0iR (7)

The defined objective function (OF) minimizes

the mass flow rate of generated VHP steam, mVHP,

defined by Eq. (1). The heat balance for one tem-

perature interval, k, is described in Figure 4.

Figure 4. Heat flows during one temperature interval.

Each temperature interval k has two inputs, heat

content from all hot stream (

,1

IHi k

i

Q ), and residual

heat from the previous temperature interval

(

, 11

I

i ki

R ). Moreover, each temperature interval has

two outputs: heat transferred to recoverable utility

(

,1

JCj k

j

Q ), and surplus heat, known as heat residual

(

,1

I

i ki

R , Eq. (2)) [39].

During heat transfer, each recoverable utility

undergoes a phase change. As a result, each rec-

overable utility has a different heat capacity value

throughout the temperature range. The heat transfer-

red to recoverable utility, Qj,k, depends on the heat

capacity at that particular temperature interval (Figure

3). Heat demands of recoverable utility, Qj,k, is equi-

valent to the sum of the transferred heat from hot

streams,

, ,1

I

i j ki

Q , to particular recoverable utility, j,

which is defined as one of the constrains in Eq. (3).

After generating recoverable utilities, j, (mass

flow rate, mj), they are set to replace the specific

required hot utility either totally or partially, req (mass

flow rate, mreq) (Figure 2). Therefore, mass flow rate

of recoverable utility, mj must be less than or equal to

the mass flow rate of corresponding required utilities,

mreq. This condition is set as one of constrains, and it

is defined by Eq. (4). All required hot utilities, req, are

supplied by VHP steam, which is generated in com-

mon utility system using fossil fuels. Hence, gener-

ated recoverable utilities, j, replace the required hot

utilities, req and consequently reduce demands for

generated VHP steam, mVHP, in common utility (Eq. (5)).

Known parameters are set of heat sources,

,Hi kQ , i Hk , at each temperature interval and the

mass flow of each required hot utility, mreq:

, , 1( )Hi k p i k kQ C , ,i Hk k TI (8)

, ,

, , ,

/ ( ( )

( )),

req req vap req sat req

pG req uh req sat req

m Q H

c req RHU (9)

Optimized variables are mass flow rates of rec-

overable utilities, mj, j Ck , which define the heat

content, ,Cj kQ , represented in Eq. (10):

, , 1( )Cj k j p j k k kQ m c , ,j Ck k TI (10)

As aforementioned, recoverable utilities undergo

a phase change during the heat transfer therefore

phase state and specific heat capacity depends on

the specified temperature interval. This is represented

in Eq. (11) by three options for each of the phase

state: liquid phase (L), vaporization (VAP) and super-

heated state (SS):

,

, ,

,

,

, 1

, 1

L

VAP

SS

in satp j k k

sat satp j k p j k k

sat outp j k k

c T T

c c T T

c T T

(11)

,j Ck k TI

where

, / Tp j k vap jc H , 1T K (12)

During the vaporization, saturation temperature,

Tsat, is constant, but in this model formulation it is

approximated that the vaporization is happening

throughout 1 K. The values are obtained from the

standard thermodynamic tables [40]. Each specific

heat capacity value, cp, is calculated as an average

value of the two values: at the initial and final tempe-

rature of the phase state.


78

Case study

The proposed model has been applied to an

industrial case study [9]. An industrial zone consists

of four independent petrochemical plants, where each

plant consists of one or more processes. Each plant is

served by an independent utility system and each

plant has been optimized for energy efficiency. The

methodology to develop optimal heat recovery in

decentralized system is presented in previous study

[9]. In this study, the industrial zone is considered to

be served by one centralized utility system, which

provides required heat for all processes, instead of

each plant having independent utility system.

The fossil fuel used in the utility system is nat-

ural gas. Combustion of natural gas generates VHP

steam, which is expanded by a let-down station to

lower pressure and temperature: HP, MP or LP

steams, in order to satisfy the requirements set by

processes in each plant.

Data acquisition

For required optimization, three sets of data are

necessary. The first set of data represents hot pro-

cess streams and excess heat that can be reused.

The hot streams are defined by the Stijepovic and

Linke method using the concept of exergy is applied

[9]. There are seven hot process streams that are

recognized as potential heat sources and their pro-

perties are summarized in Table 1.

Table 1. Data for hot process streams

Stream number Plant number Tin / C Tout / °C H / kW

1 1 230 60 30000

2 1 200 55 20000

3 1 55 40 10000

4 2 200 60 20000

5 3 330 60 25000

6 3 300 70 20000

7 4 180 60 12000

The second set of data represents required

utility usage in each plant. The third set of data rep-

resents properties for recoverable utilities, which are

generated in common utility via heat exchange with

hot process streams. For each of the nine required

utilities, nine recoverable utilities are introduced in the

system. The second and third set of data are sum-

marized in Table 2. All represented utilities only use

steam at different levels.

The approach assumes that all process plants in

an industrial zone are linked through the central utility

system. Hot process streams (Table 1) generate rec-

overable utilities (Table 2) in a common utility. They

are then transported to a specific plant to replace the

specified required utilities either totally or partially

(Table 2).

In order for recoverable utility to replace the

required utility, conditions like temperature and pres-

sure must match. Pressure of recoverable utility must

be the same as the pressure of required utility and the

target temperature, i.e., outlet temperature, of rec-

overable utility must match the temperature of

required utility header. In order for heat transfer to be

feasible the minimum allowable temperature differ-

ence, Tmin, must be introduced for recoverable

utilities: 30 C for HP and MP steams, and 15 C for

LP steams.

For illustration purposes, the data used are

imaginary but comparable to data observed in exist-

ing production plants.


Equations (1)-(12) form linear problem that is

solved using LINGO software [41]. The objective

function is to minimize VHP steam generated in the

centralized utility system through waste heat recovery

and reuse via indirect heat integration. Variables that

are optimized are mass flow rates of recoverable util-

ities.

Table 2. Data for required utilities

Str. No. Plant No. Required utilities Recoverable utilities

Tin / C Tout / C (Tsat) Heat required, kW Mass flow rate, kg/s Tsat / C Tin / C Tout / C

1 1 280 240 23000 12.09 240 108 280

2 1 240 200 29000 14.14 200 108 240

3 2 320 240 20000 9.88 240 108 320

4 2 250 220 39000 20.02 220 108 250

5 2 200 170 21000 9.89 170 108 200

6 2 150 120 28000 12.37 120 108 150

7 4 220 190 19000 9.23 190 108 220

8 4 150 130 25000 11.27 130 108 150

9 4 120 108 23000 10.18 108 108 120


79

Figure 5 represents indirect heat integration for

this case study. As the depicted utility system burns

natural gas, it generates VHP steam, which is red-

uced to required HP steam, MP steam and LP steam.

Recoverable utilities are generated through heat

exchange from hot process streams, as it is shown in

Figure 5. Generated recoverable utilities obtained by

optimization are four utilities with lowest saturated

temperatures: utilities 5, 6, 8 and 9. Recoverable util-

ities are directed to steam headers, where they are

linked to the corresponding required utilities. The rep-

lacement of required utilities decreases demands of

total VHP steam generated in utility system, as well

as decreases consumption of natural gas.

The obtained results of optimized variables,

mass flow rate of recoverable utilities, mj, are pre-

sented in the Table 3. Results of the optimization

show that all generated recoverable utilities are low

Figure 5. Case study system with indirect heat integration.

Table 3. Comparing mass flow rates of required utilities before and after heat integration

Utility Tsat / C Tout / C Mass flow rate of recoverable utility,

mj / kg s-1

Mass flow rate of required utility before

heat integration, kg/s

Mass flow rate of required utility

after heat integration, kg/s

1 240 280 0 12.09 12.09

2 200 240 0 14.14 14.14

3 240 320 0 9.88 9.88

4 220 250 0 20.02 20.02

5 170 200 1.16 9.89 8.73

6 120 150 12.37 12.37 0

7 190 220 0 9.23 9.23

8 130 150 11.27 11.27 0

9 108 120 10.18 10.18 0

Total – – 34.98 109.07 74.09


80

grade utilities: utilities 5, 6, 8 and 9. Three out of four

generated utilities are totally generated with targeted

mass flow rate. These results are expected, since the

objective function is set to minimize the quantity of

used VHP steam as heat resource.

Additionally, in Table 3 are represented data for

the mass flow rates of required utilities before heat

integration and after heat integration. The mass flow

rate of each required utility after heat integration rep-

resents difference between mass flow rate of required

utility before heat integration and mass flow rate of

corresponding recoverable utility, which is set to rep-

lace required utility. The requested total mass flow

rate by required utilities before the heat integration is

109.07 kg/s. After the heat integration, total mass flow

rate is reduced to 74.09 kg/s, which is 32.07% less

than before the integration. Therefore, demands for

generating VHP steam in utility system is reduced by

32.07%.

Heat determined by required utilities supplied by

the utility system, before and after heat integration is

presented in Table 4. As it can be observed, the opti-

mization results show that 78.463 kW heat can be

recovered via heat integration, which represents

34.56% of heat supplied by steam from centralized

utility system.

Less demands for generating VHP steam, after

the heat integration, leads to less demands for

combustion of natural gas in the centralized utility

system. The consumed natural gas is compared

before and after heat integration in order to evaluate

the amount of natural gas that can be saved through

heat integration in this case study. Natural gas

consumption is defined by next equation:

VHP , , VAP

, ng ng

( ( )

( ))

pL j sat j in

pG j out sat

m c H

c m LHV (13)

where mVHP is total mass flow rate of VHP steam that

is generated from burning natural gas (kg s-1), mng

mass of natural gas (kg s-1), and LHV is low heat

value for natural gas (kJ kg-1). Comparing the amount

of natural gas consumed before (7.89 kg s-1) and after

heat integration (5.36 kg s-1), the saved amount of

natural gas during the heat integration is 2.53 kg s-1.

CONCLUSION

A method for waste heat recovery and reuse via

heat integration is developed in order to increase

energy efficiency and decrease the use of fossil fuels.

Industrial zones usually consist of multiple independ-

ent plants, where each plant is supplied by an inde-

pendent utility system, as a decentralized system. In

this study, a method is applied to target minimum

energy requirements where an industrial zone is sup-

plied by a centralized utility system instead of decen-

tralized utility system. The proposed method is based

on linear programming problem (LP). It was tested out

on a case study and the results indicate that fossil fuel

savings are achieved, and energy efficiency of an

industrial zone is increased by recovering and reusing

waste heat via indirect heat integration. This approach

can be used in the decision making process in retro-

fitting strategy for utility system configuration in an

industrial zone.

Acknowledgments

This work was supported by Ministry of Edu-

cation, Science and Technology Development, Rep-

ublic of Serbia Project no. OI172063.

Nomenclature

Indices

i - hot stream

j - recoverable utility

k - temperature interval

req - required hot utility

pc - phase change of recoverable utility

Sets

Hk = {i | hot stream i supplies heat at interval k, i =

= 1,…,I}

Ck = {j | recoverable utility j demands heat at interval

k, j = 1,...,J}

TI = {k | temperature interval, k = 1,…,K}

RHU = {req | required hot utility, req = 1,…,REQ}

Table 4. Comparing data of required utilities before and after heat integration, heat, kW

Utility No. Heat required before the heat integration Heat required after the heat integration Heat recovered via heat integration

5 21000 18536 2464

6 28000 0 28000

8 25000 0 25000

9 23000 0 23000


81

Parameters

,Hi kQ - heat content of hot stream i at temperature

interval k, kW

reqQ - heat content of required hot utilities req, kW

mreq - mass flow rate for required utility, kgs-1

Cp,i - heat capacity flowrate of the hot process

streams, i, kJ K-1

cpG,req - specific heat capacity of the gas phase, G, for

required utility, req, kJkg-1 K-1

Hvap,req - latent heat of vaporization for required utility

stream, req, kJ kg-1

θuh,req, θsat,req - temperature of utility header and sat-

uration temperature of required utility, K

Variables

mj - mass flow rate for recoverable utility, kg s-1

,Cj kQ - heat content demanded by recoverable utility j

at temperature interval k, kW

, ,i j kQ - heat exchanged between hot stream i and

recoverable utility j at temperature interval k, kW

,i kR - heat residual of hot stream i at temperature

interval k, kW

θk , θk-1 - inlet and outlet temperatures of the each

temperature interval, k, K

cp,j,k, - specific heat capacity of recoverable utility j, kJ

kg-1 K-1

mVHP - mass flow rate of generated VHP steam, kg s-1

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MILANA M. ZARIĆ1

MIRKO STIJEPOVIC1

PATRICK LINKE2

JASNA STAJIĆ-TROŠIĆ1

BRANKO BUGARSKI3

MIRJANA KIJEVČANIN3

1Institut za hemiju, tehnologiju i

metalurgiju, Univerzitet u Beogradu,

Njegoševa 12, 11000 Beograd, Srbija 2Department of Chemical Engineering,

Texas A&M University at Qatar, P.O.

Box 23874, Education City, Doha,

Qatar 3Tehnološko-metalurški fakultet,

Univerzitet u Beogradu, Karnegijeva 4,

11120 Beograd, Srbija

NAUČNI RAD

REKUPERACIJA TOPLOTE U INDUSTRIJSKOJ ZONI

Sa ciljem da se smanji upotreba fosilnih goriva u industrijksim sektorima, a da zahtevi pro-

cesa proizvodnje budu zadovoljeni, razvijaju se novi pristupi toplotne integracije i reku-

peracije toplote. Cilj ove studije je razvijanje pristupa koji će omogućiti povećanje energet-

ske efikasnosti u industrijskim zonama rekeperacijom otpadne toplote putem indirektne

integracije. Uobičajeno je da se industrijska zona sastoji od više nezavisnih postrojenja,

kao decentralizovani sistem, gde je svako postrojenje obezbeđen nezavisinim sistemom

pomoćnih fluida. U ovoj studiji, razvijen je novi pristup, gde se minimalizuju energetski zah-

tevi i gde se industrijska zona obezbeđuje centralizovanim sistemom pomoćnih fluida,

umesto decentralizovanog sistema. Ovaj pristup pretpostavlja da su sva procesna postro-

jenja u industrijskoj zoni povezana kroz centralizovani sistem pomoćnih fluida. Predloženi

metod je formulisan kao problem linearnog programiranja (LP). Pored toga, ovaj postupak

može se koristiti tokom odlučivanja o strategiji energetske integracije industrijskih zona.

Štaviše, predloženi metod je primenjen na studiju slučaja. Rezultati pokazuju da je moguće

ostvarenje uštede fosilnih goriva.

Ključne reči: rekuperacija toplote, energetska efikasnost, toplotna integracija, LP

programiranje.






83

SANDRA RAQUEL KUNST1

LILIAN VANESSA ROSSA

BELTRAMI2

MARIELEN LONGHY1

HENRIQUE RIBEIRO

PIAGGIO CARDOSO2

TIAGO LEMOS MENEZES2

CÉLIA de FRAGA MALFATTI2

1Programa de Pós Graduação em

Engenharia de Processos e

Tecnologias (PGEPROTEC),

University of Caxias do Sul (UCS),

Caxias do Sul, RS – Brasil 2Laboratório de Pesquisa em

Corrosão (LAPEC), Federal

University of Rio Grande do Sul

(UFRGS), Porto Alegre, RS - Brasil

SCIENTIFIC PAPER

UDC 669.715:678:544:543.42


EFFECT OF DIISODECYL ADIPATE CONCENTRATION IN HYBRID FILMS APPLIED TO TINPLATE

Article Highlights

• Tinplate protection from corrosion

• Flexible hybrid film obtained by sol-gel process

• Electrochemical impedance spectroscopy tests

Abstract

Siloxane hybrid films are fragile and have low mechanical strength due to their

vitreous material properties. Hence, a new formulation incorporating a plasticizer

agent was developed in order to increase the layer thickness of a uniform and

homogeneous hybrid film on tinplate, and to provide flexibility to the polymeric

matrix. Tinplate sheets were coated with a hybrid film obtained from a sol-gel

process, constituted by t+he addition of the following alkoxide precursors:

3-(trimethoxysilyl) propyl methacrylate and tetraethoxysilane with 0.01 mol L-1

cerium nitrate addition. The influence of the diisodecyl adipate plasticizer con-

centration was evaluated. The films were characterized by scanning electron

microscopy, profilometry, open circuit potential monitoring, polarization curves

and electrochemical impedance spectroscopy. The results showed that all films

with diisodecyl adipate had higher electrochemical performance compared to

uncoated tinplate. However, the film with the 2% plasticizer concentrations had

the best performance in the electrochemical tests, although it had thinner layer.

Keywords: hybrid film, diisodecyl adipate, tinplate, electrochemical behaviour.

The sol-gel process is a widely researched and

used technology [1-5], as a result of the obtained

surface protection properties, the simplicity of the

formulation process and its economic viability. Among

its advantages, the following can be mentioned: the

stoichiometry is easy to control and adjust [6]; it is

possible to tailor a film with high purity and with a uni-

form distribution of the components [7]; the process

can be carried out under normal pressure and low

temperatures [8-12]. Furthermore, the sol-gel process

has been used for decades to obtain a high number of

hybrid materials, using inorganic and polymeric pre-

cursors [13,14].

Correspondence: S.R. Kunst, Programa de Pós Graduação em

Engenharia de Processos e Tecnologias (PGEPROTEC), Uni-

versity of Caxias do Sul (UCS), Caxias do Sul, RS – Brasil. E-mail: [email protected] Paper received: 25 July, 2015 Paper revised: 6 November, 2015 Paper accepted: 11 March, 2016

The sol-gel process consists of the hydrolysis

and condensation of alkoxide precursors with metal-

oxane, in order to obtain a three-dimensional siloxane

network. After deposition of the film on the substrate

by means of appropriate techniques, the film is

exposed to air at the beginning of the condensation

reaction. After a few minutes of drying, a network (gel)

is formed on the substrate [15,16]. The formed net-

work is a hybrid (organic and inorganic) [16]. Water

molecules are eliminated through sintering (densific-

ation) at an appropriate temperature, and a compact

hybrid layer is formed. Then, the film is submitted to a

final thermal treatment for structure control [17].

Hybrid films are used to coat a variety of metallic

substrates to protect them from corrosion [18,19].

Amongst these, tinplate can be mentioned. Tinplates

are used in food and non-food packaging applications

[20]. For good quality, the presence of tin oxides on

the surface and the condition of the passivation layer

are very relevant. The presence of such oxides can

S.R. KUNST et al.: EFFECT OF DIISODECYL ADIPATE CONCENTRATION… Chem. Ind. Chem. Eng. Q. 23 (1) 8395 (2017)

84

alter the sheet appearance, weldability and capability

of being coated with organic films [21]. Currently,

many surface treatments for packaging are based on

chromates, as they offer excellent corrosion resist-

ance, but their application process is highly toxic.

Thus, non-toxic pre-treatment alternatives have been

developed recently to replace the chromating process

[22,23]. Hybrid films obtained by the sol-gel process

are one of those alternatives and the industrialization

of the sol-gel process involves work with real sub-

strates [16,24-26].

Research has been performed on the tinplate

coating with hybrid films. Kunst et al. [27] studied the

coating the tinplate with mono and bilayer of hybrid

films modified with polyethylene glycol (PEG). These

hybrid films were obtained from precursors 3-(trimeth-

oxysilyl)propylmethacrylate (TMSM) and tetraethoxy-

silane (TEOS) with the addition of cerium nitrate in a

concentration of 0.01 M. The hybrid films were

applied in single and double layers and cured at dif-

ferent temperatures (60 and 90 C). The hybrid film

with monolayer and cured at 90 C showed a more

compact, uniform, less porous layer and better elec-

trochemical impedance results.

Kunst et al. [28] studied the effect of the con-

centration of tetraethoxysilane (TEOS) on the pro-

tective properties of the film on tinplate substrate. The

tinplate was coated with a hybrid film obtained from a

sol-gel method constituted of the alkoxide precursors

3-(trimethoxysilyl)propylmethacrylate (TMSM) and

poly(methyl methacrylate) (PMMA) and different con-

centrations of tetraethoxysilane (TEOS). The hydro-

lysis of these films was performed at a pH value of 3

using acetic acid as a catalyst. All the studied films

have shown good performance as to corrosion resist-

ance on tinplate, but the film with ratio TEOS:MPTS

of 1 showed the best electrochemical results.

Kunst et al. [29] studied protective coatings as

hybrid films composed by different acids are studied

to improve the barrier effect against corrosion. The

hybrid films deposited on a tinplate from a sol made

up of the alkoxide precursors 3-(trimethoxysilyl)pro-

pylmethacrylate (TMSM), tetraethoxysilane (TEOS)

and poly(methylmethacrylate) (PMMA) with benzoyl

peroxide (BPO). The hybrid sols were prepared by

mixing water with three different acids: acetic

(CH3COOH), hydrochloric (HCl) and nitric acid (HNO3).

The results demonstrate that the hybrid film obtained

by acetic acid addition exhibiting the greatest

improvement the protective properties of the tinplate.

Malfatti et al. [30] showed that the addition of

Ce3+ in a hybrid films confers additional active cor-

rosion protection through a self-healing behavior that

can be verified by the reduction in corrosion rate com-

pared to hybrid films without Ce3+ addition. Thus,

considering the industrial potential for application of

the coating, it is expected that the Ce3+ should dec-

rease the corrosion process in case of damage to the

barrier layer.

In order to improve the barrier effect and coat

the tinplate uniformly and homogeneously, it is desir-

able to obtain coatings with higher layer thickness.

For this purpose, there are two possible approaches:

one is to increase the number of layers up to a limit

number to avoid delamination [31]. An alternative

approach is to increase the sol viscosity; this para-

meter can be modified either by varying the tempe-

rature to alter the hydrolysis and condensation react-

ion kinetics during the siloxane film formation, or by

introducing a plasticizer agent to modify the intrinsic

properties of the gel. The latter was chosen in this

study, due to the fact that it is easily adjustable [15].

The purpose of this work was to coat tinplate

with hybrid film, obtained from a sol-gel process. This

film was constituted by the alkoxide precursors 3-(tri-

methoxysilyl)propyl methacrylate (TMSM) and tetra-

ethoxysilane (TEOS) and was added different diiso-

decyl adipate plasticizer concentrations (0.5, 1, 2 and

4%). The films were applied using a by dip-coating

process and were cured at 60 C for 20 min. From

this study, an ideal concentration of diisodecyl adipate

plasticizer was determined for the training of a hybrid

film with protective properties superior to other films

studied.


Surface preparation

Tinplate coupons with dimensions of 20 mm40

mm were obtained from industrial sheet, rinsed with

acetone and dried. The tinplate used in this study has

the maximum percentage composition of: 0.06C,

0.2Mn, 0.02P, 0.02S, 0.005N and 0.06Al. The aver-

age thickness of the tinplate used was 0.245 mm. The

used tinplate had yield strength ranging from 210 to

310 MPa, tensile strength of 290 to 410 MPa, hard-

ness of 51 to 59 HR 30 T D and tin coating of 3.0 to

2.0 g m-2. The plate samples were degreased with a

10 min immersion in neutral detergent at 70 C,

washed and dried.

Hybrid films elaboration

Hydrolysis reactions were conducted with the

silane precursors 3-(trimethoxysilyl)propylmethac-

rylate (TMSM, C10H20SiO5) and tetraethoxysilane

(TEOS, C8H20SiO4), with 0.01 mol L-1 cerium nitrate


85

addition, and water and ethanol were used as sol-

vents. The TEOS/TMSM/H2O/EtOH proportions were,

respectively, 25/4/16/55 wt.%. A diisodecyl adipate

plasticizer was added to the sol formulation in four dif-

ferent concentrations (0.5, 1, 2 and 4 wt.%); a sample

without plasticizer addition also was analyzed. The

hydrolysis time was of 24 h at 25 C. A dip-coating

process was conducted to coat the substrate with the

hydrolyzed hybrid solution, with a removal rate of 10

cm min-1 and a residence time of 5 min. After the dip-

-coating process, the hybrid film pre-treated sub-

strates were thermally oven-cured at 60±2 C for 20

min.

Table 1 presents the description of the studied

samples, and the entire protocol of the sol preparation

and coatings is detailed in the flow chart presented in

Figure 1.

Experimental techniques

Morphological characterization was performed

using a JEOL 6060 scanning electron microscope

(SEM) at an acceleration voltage of 20 kV. The

samples were observed from top surface view and in

a cross-section to determine the layer thickness. The

surface micro-roughness was evaluated in a contact

profilometer (PRO500 3D). The wettability of the hyb-

rid films was evaluated by contact angle measure-

ments through the sessile drop method in equipment

developed by the Laboratory of Corrosion Research

(LAPEC) at UFRGS. The contact angle was deter-

mined using image analysis software.

The corrosion performance of the coatings was

evaluated by open circuit potential (OCP) monitoring,

polarization curves and electrochemical impedance

spectroscopy (EIS) measurements in a 0.05 mol L-1

NaCl solution. All electrochemical tests were per-

formed in triplicate, in an environment temperature

and without aeration. This concentration is sufficiently

high to activate corrosion in a relatively short expo-

sure time but is low enough to enable the effects of

the plasticizer to be determined. Kozhukharov et al.

[32] also used 0.05 mol L-1 NaCl to ensure a suf-

ficiently low concentration to allow for the observation

of corrosion inhibitor effects. A three-electrode cell

was used to perform the evaluations, with a platinum

wire as the counter-electrode and a saturated calomel

electrode (SCE) as the reference electrode. The area

of the working electrode was 0.626 cm2. The polariz-

ation curves were collected at a scan rate of 1 mV s-1

and potential intervals between 200 mV (below OCP)

and 400 mV (above OCP).

From the extrapolation of Tafel slopes at the

polarization curves were determined the corrosion

potential (Ecorr), corrosion current (icorr), polarization

resistance (Rp) and the protection efficiency (PE) for

samples studied. PE of the coatings was determined

for the equation, where icorr and corr*i are the corrosion

current densities obtained for uncoated and coated

substrates, respectively:

Table 1. Description of the samples

Sample Description

F1A60M Tinplate coated with hybrid film with 0.5% diisodecyl adipate plasticizer addition

F2A60M Tinplate coated with hybrid film with 1% diisodecyl adipate plasticizer addition



F5A60M Tinplate coated with hybrid film without plasticizer addition

Fl Uncoated tinplate

Figure 1. Schematic representation of coating elaboration.


86

corr corr

corr

* 1 00

i iPE

i

For the EIS measurements, the samples were

available up to 96 h and the systems were monitored

by open circuit potential (OCP) for 1 h prior to each

test. The amplitude of the EIS perturbation signal was

a sinusoidal 10 mV (rms) signal, and the frequency

range was from 100 kHz to 10 mHz using a Nova®

frequency response analyzer and a Autolab PGSTAT

30 potentiostat. The results obtained were fitted using

electrical equivalent circuits (EEC) using the Nova®

software. The consistency of the experimental data

was verified with the Kramers-Kronig transform

(KKT), and data that did not match were discarded.


Morphological characterization

SEM micrographs were performed for the hybrid

films with plasticizer addition F1A60M (0.5% plas-

ticizer), F2A60M (1% plasticizer), F3A60M (2% plas-

ticizer) and F4A60M (4% plasticizer), and the hybrid

film without plasticizer addition F5A60M before the

electrochemical tests. They analyses are shown in

Figure 2.

The micrographs show that the hybrid films with

less diisodecyl adipate plasticizer concentration

F1A60M and F2A60M (0.5 and 1%) were the only

samples to show cracks and the F4A60M sample

shows discontinuities in the film surface.

The presence of cracks in the films with a low

concentration of plasticizer was due to silane pre-

cursors TMSM and TEOS that tend to form a compact

tridimensional network and, due the presence of the

plasticizer, the latter is encapsulated in the network

due to the interaction forces (i.e., hydrogen bonds,

van der Waals forces or covalent bonds). Thus,

branch mobility was limited, due to the absorption of

mechanical energy. This produces a more rigid and

fragile hybrid film. The phenomenon is known as anti-

plasticizing [33]. The presence of discontinuities in the

film with the highest concentration of plasticizer can

be associated with incompatibilities between sol and

plasticizer. According to Grossman et al. [34], adipate

plasticizers typically are used from C7 to C10 and

heterogeneities can rise when the plasticizer is added

in higher concentrations for a higher number of car-

bons. As TEOS is a C8 and TMSM is a C10 molecule,

this explanation accounts for the discontinuities obs-

erved in Figure 2 (F4A60M).

The increase in thickness of the films with the

addition of diisodecyl adipate could be due the plas-

ticizer, which enters into polymeric chains, thus inc-

reasing their free volume. Adipates are external plas-

ticizers; i.e., they are additives that interact physically

with the polymer. There can be weak attractive force

between the polymeric matrix of the film and the plas-

ticizer, through weak force as hydrogen bonds and/or

van der Waals forces, though adipates do not react

chemically with the silane precursor radical. Due to its

smaller molecular size compared with the polymeric

matrix of the film, diisodecyl adipate promotes an

increase in mobility of the film, flexibility of the mole-

cules, attributed to the generated increase of free vol-

ume, enabling an increase in the layer thickness for

those systems.

The thickness of the hybrid film layer was

determined from a cross-section SEM image analysis

(Figure 2) and these results are shown in Table 2. As

can be observed, all hybrid films with diisodecyl adi-

pate addition (F1A60M, F2A60M, F3A60M and

F4A60M) showed a significant layer thickness inc-

rease, compared to the sample without plasticizer

addition (F5A60M). The F1A60M and F2A60M hybrid

films had the higher layer thickness values; however,

they also exhibited cracks (Figure 2).

However in this study, it was observed that the

hybrid films with low plasticizer concentration pre-

sented higher layer thickness values, but also exhi-

bited cracks (Figure 2). This was due to the fact that,

despite of the adipate plasticizers being external plas-

ticizers, i.e., that they only interact physically with the

polymer; this additive is esterificated by the adipic

acid after esterification, and hence they pertain a

reactive group. A problem that can rise with such

reactive groups is that they can react with the poly-

meric matrix of the coating, making the film molecular

size bigger, thereby reducing its flexibility [35], and

promoting crack formation.

The results of the contact angle measurements

for the different systems studied are shown in Table

2. According to the results, all of the hybrid films

showed hydrophilic behavior and there were no sig-

nificant differences among the contact angle values

for the systems. The samples were hydrophilic in nat-

ure and there was not a significant difference among

the contact angle values for the systems, because the

diisodecyl adipate plasticizer is not involved in the

silane hydrolysis and condensation reactions. The

predominant interactions between the plasticizer and

the polymeric matrix are hydrogen bonds or van der

Waals forces, which are weak interactions and do not

interfere in film formation.

Silane precursor-based hybrid films show hydro-

phobic behavior when they are not sufficiently cross-


87

Figure 2. SEM images and layer thickness of the hybrid films.


88

Table 2. Layer thickness, contact angle and surface roughness values for the hybrid films studied

Sample Layer thickness, µm Contact angle, Surface roughness

Ra / µm Rms / µm Peak-to-peak, µm

F1A60M 3.93±0.31 72±1.09 0.37±0.09 0.47±0.08 4.41±0.16

F2A60M 3.21±0.36 77±0.72 0.36±0.11 0.48±0.10 3.91±0.13

F3A60M 1.96±0.18 76±0.87 0.31±0.08 0.39±0.09 4.11±0.11

F4A60M 2.14±0.24 66±0.60 0.34±0.09 0.43±0.08 2.57±0.09

F5A60M 0.63±0.11 78±1.50 0.41±0.12 0.50±0.11 3.53±0.12

Fl – 73±1.90 0.43±0.10 0.51±0.10 2.40±0.11

-linked. During the curing process, siloxane bonds

form a network that hinders water penetration, con-

sequently enhancing the hydrophobic character.

Hence, the contact angle for a highly cross-linked film

is about 90. Neither silane precursors hydrolysis nor

cross-linking (polycondensation) are completed during

curing at 60 C [36]. Subsequently, therefore, non-

-hydrolyzed ester and hydrophilic –OH groups are pre-

sent in the hybrid films structures and favor water

absorption [37].

All of the studied hybrid film samples contained

Ce in their formulation and in accordance with a lite-

rature survey performed by Palomino et al. [38], the

results suggest that this behavior was due to the joint

action of two factors: i) improved barrier properties

imparted to the silane film due to the incorporation of

silica particles, which would block preferential path-

ways for electrolyte penetration, ii) the higher degree

of polymerization and increased thickness of the sil-

ane layers, due to the presence of cerium ions. The

direct influence of Ce can be observed in the results

of contact angle for F5A60M sample (without plasti-

cizer) where the presence of Ce promoted the form-

ation of a more hydrophobic film than was the case

with the other films.

Roughness values are summarized in Table 2,

where Ra stands for arithmetic average, Rms for aver-

age square roughness, and Ry for maximum rough-

ness or peak to trough size. There was no evident

consistent variation in the roughness values, due to

the low resolution capability of the technique (i.e., a

variation of only micrometers). However, all of the

hybrid films with plasticizer addition showed better

roughness values compared to the F5A60M sample

(without adipate) and to uncoated tinplate.

Electrochemical characterization

Open circuit potential (OCP) monitoring was

conducted in a 0.05 mol L-1 NaCl solution to verify the

potential variation with time. These results are shown

in Figure 3a. Also, polarization curves for all the sys-

tems are presented in Figure 3b. From Tafel slope

extrapolations (Table 3), the corrosion current density

(icorr), corrosion potential (Ecorr) and polarization resist-

ance (Rp) values were determined.

Open circuit potential values (Figure 3a) for 1 h

of immersion indicate that all hybrid films had poten-

tials shifted upwards in relation to the uncoated tin-

plate (Fl), in the following order: F1A60M (-390 mV) >

F4A60M (-440 mV) > F5A60M (-448 mV) > F2A60M

(-450 mV) > F3A60M (-468 mV) > Fl (-509 mV); that

is, the hybrid films presented less active potentials

and promoted the formation of a barrier between the

substrate and the medium.

Furthermore, the potential for the F1A60M sys-

tem shifted in the positive direction (–390 mV) com-

Figure 3. a) Open circuit potential and b) polarization curves in a 0.05 mol L-1 NaCl solution for all studied hybrid films

and for tinplate substrate.


89

pared to the uncoated tinplate sample Fl (–509 mV);

i.e., this difference in potential value results from a

barrier effect from the hybrid film [39]. Moreover, the

equilibrium potential that was observed for the tinplate

substrate after 1 h immersion (–509 mV vs. SCE) is

associated with the formation of a Sn(II)-oxide/hydro-

xide layer [40,41].

With regard to the polarization curves (Figure

3b), it is observed that all of the hybrid films (F1A60M,

F2A60M, F3A60M, F4A60M and F5A60M) showed a

decrease (two orders of magnitude) in the corrosion

current density (icorr) in comparison with the uncoated

tinplate sample, which denotes the protective action

of those films. The polarization curves also showed

an increase by almost one or two orders of magnitude

in terms of the polarization resistance value (Table 3).

Analyzing the hybrid films only, the siloxane-adi-

pate film obtained with a 2% concentration of diiso-

decyl adipate, which corresponds to the sample

F3A60M (icorr = 0.914 mA m-2 and Rp = 73.2 Ω m2)

showed the best performance of the studied hybrid

films, yet only marginally so.

From the polarization curves it can be observed

that the tinplate substrate showed higher values of

current density (icorr = 47.1 mA m-2) and a lower pola-

rization resistance (Rp = 0.554 m2) compared to the

studied hybrid films.

From polarization tests showed that all of the

hybrid films exhibited better protective performance

than the sample without a hybrid film. This occurs due

to the covalent bonds of the organic and inorganic

precursors, demonstrating the synergistic effect of the

TMSM and TEOS precursors’ presence in the hybrid

film.

Of the hybrid films, the sample with 2% plas-

ticizer performed better, due to the presence of weak

bond strength between the plasticizer and the silane

precursors radicals and the increased layer thickness.

However, the film was more rigid and fragile [35].

The poor performance of uncoated tinplate

could be related to the presence of tin oxide/hydro-

xide that partially passivates the metallic surface. The

passivity breakdown at –430 mV vs. SCE is related to

pitting attack that is induced by the chloride effect

[42]. Galic [43] suggests that the adsorption of chlo-

ride ions at the oxide-electrolyte interface leads to the

formation of a tin-oxychloride film, with less protective

behavior.

Nyquist and Bode plots obtained by electroche-

mical impedance spectroscopy tests for 24 and 96 h

of immersion in 0.05 mol L-1 NaCl solution for all the

studied systems and for tinplate are shown in Figure

4. Table 4 shows the RHF (resistance for high fre-

quency) and RCP (resistance caused by corrosion pro-

cess).

The impedance results obtained from the

Nyquist plot (Figure 4) show higher resistance values

for the sample F3A60M (RHF = 472 k cm2) after 24 h

of immersion. This is due to the fact that these coat-

ings had no cracks or discontinuities, as was obs-

erved on the SEM images (Figure 2). However, after

96 h of immersion the resistance for the system

F3A60M (RHF = 312 k cm2), the resistance dim-

inished by a factor of two. Furthermore, the F3A60M

sample showed a higher phase angle (around 68 -

Figure 4) and a higher impedance modulus (around

log |Z| = 5.25 - Figure 4) values, after 96 h of immer-

sion.

On the other hand, the F1A60M sample, which

had a high layer thickness and had the addition of

0.05% plasticizer showed lower resistance values

after 24 (RHF = 92.71 k cm2) and 96 h (RHF = 4.31 k

cm2) of immersion. These reveal the fragility of this

coating, with cracks that allow permeation of the elec-

trolyte through the film and towards the substrate.

The EIS tests results indicated that the good

electrochemical performance with 2% of diisodecyl

adipate plasticizer occurs because the plasticizer con-

centration is the most suitable to increase the free

volume, allowing the movement of the polymeric chain

and giving flexibility to the organic radicals bound to

the silicon atom. Also, plasticizer inclusion avoids

interactions between the organic radicals in the silane

precursors, enhances the mobility of the chain ext-

remities and improves the flexibility of the system.

Table 3. Obtained data from Tafel extrapolation

Sample icorr / mA m-2 Ecorr / mV Rp / m² PE / %

F1A60M 0.566 -417 72.5 98.80

F2A60M 0.574 -464 56.0 98.78

F3A60M 0.914 -504 73.2 98.06

F4A60M 0.778 -451 42.1 98.35

F5A60M 0.570 -497 68.0 98.79

Fl 47.1 -581 0.554 -


90

EIS tests also revealed the poor performance of

the hybrid film with less plasticizer concentration.

Although the formulation of these systems has pro-

moted an increase in the layer thickness (Figure 2

and Table 2) due to the plasticizer addition, a weak

and porous structure resulted, due to the formation of

intertwined molecules liked only weak bonds (hydro-

gen bonds). Moreover, it restricted the mobility of

small branches in the chain, causing cracks on the

films, which contribute to the poor corrosion resist-

ance of these samples and their consequent inability

to resist long periods of immersion.

It was observed that after 96 h of immersion the

hybrid films showed a decrease in their protective

action, which reduced their efficiency in the corrosive

environment. This behavior was due the permeability

of the film, which occurred during the test and red-

uced the barrier properties of the film [44].

Figure 4. Nyquist and Bode plots obtained for hybrid film coated and uncoated tinplate in NaCl 0.05 mol L-1 solution for:

a) 24 and b) 96 h of immersion.


91

Equivalent circuit models can be used to explain

the electrochemical impedance data obtained from

the EIS tests. These models, which use a com-

bination of resistance, capacitance and other elec-

trical elements to simulate the coating response, can

be helpful in providing a clear understanding of the

response of the electrochemical system [45]. In this

work, two equivalent electrical circuit models were

used (Figure 5). In several circuits, the capacitance

was substituted by a CPE in order to accommodate

the non-ideality of the systems. In these circuits (Fig-

ure 5a), Re represents the electrolyte resistance, RHF

represents the film resistance and CPEHF represent

film capacitance for high frequency. The same equi-

valent circuit model (Figure 5a) is proposed for the

electrochemical behavior simulation for all immersion

times studied (1, 24, 48, 72 and 96 h) for the

F2A60M, F3A60M and F4A60M samples. This beha-

vior was observed by the other authors [46,47], indi-

cating that the hybrid films retard the corrosion pro-

cesses on the tin plate surface.

Table 4 presents the equivalent electrical para-

meter obtained by fitting the equivalent circuit from

the EIS test data, obtained for the F1A60M, F2A60M,

F3A60M, F4A60M and F5A60M hybrid films after 96

h of immersion in a 0.05 mol L-1 NaCl solution. The

percentage errors shown in brackets in Table 4 show

that the errors involved in the fitting procedure were

less than 10% (less than 5% in most cases).

The other equivalent circuit model (Figure 5b)

was proposed for the electrochemical behavior sim-

ulation at all immersion times studied (1, 24, 48, 72

and 96 h) for the F1A60M and F5A60M samples

[40,41]. This new equivalent circuit model was emp-

loyed because the other circuit equivalent (Figure 5a)

does not allow approximation between simulation and

real data. In these circuits (Figure 5b), Re represents

the electrolyte resistance, RHF represents the film

resistance and CPEHF represent film capacitance for

high frequency, RCP and CPECP represent the resist-

ance and a constant phase element indicating the

diffusion of electrolyte through the film, initiating the

Figure 5. Equivalent electrical circuits for EIS data fitting (a,b) and evolution of the hybrid films resistance for F1A60M, F2A60M,

F3A60M, F4A60M and F5A60M in a 0.05 mol L-1 NaCl solution with the immersion time (c).


92

corrosion process [17,41,48,49]. The same equivalent

circuit model (Figure 5b) was proposed for the elec-

trochemical behavior simulation at all immersion times

studied (1, 24, 48, 72 and 96 h) for the F1A60M and

F5A60M samples. This model is consistent with the

protective action lower of these systems. The system

resistance values are found to be around (RHF = 4.31

k cm2 for F1A60M and RHF = 8.51 k cm2 for

F5A60M) after the 96 h of immersion (Table 4).

Table 4 shows that the sample F3A60M had

lower capacitance values than the other analyzed

samples until the end of the test (CPE = 11.89 µF cm–2),

denoting the best protective barrier properties of the

hybrid film. Furthermore, there was an increase in the

coating capacitance of the F1A60M sample after 1

hour of immersion (CPE = 3.29 µF cm–2, Table 4) and

after 24 h of immersion (CPE = 3.69 µF cm–2, Table 4).

For the samples F2A60M, F4A60M and

F5A60M, a slight increase capacitance values (Table

4) was observed in the mean 48 h for 96 h immersion.

After 96 h of immersion, another small significant inc-

rease in the capacitance of the film was observed.

The film capacitance always tends to increase

with immersion time as result of electrolyte uptake.

This behavior is due to a significant increase of the

dielectric constant of the coating, which is influenced

strongly by electrolyte penetration into the coating.

Moreover, chi-square (2) was introduced in Table 4

and the observe values were around 10–3, similar to

results obtained for Sakai et al. (2012) [40].

Figure 5c shows the evolution of the coating

properties (i.e., the resistance) as a function of

immersion time. The F3A60M (RHF = 867 k cm2)

sample exhibited the highest resistances for 1 h

immersion. Furthermore, the F3A60M sample exhi-

bited the highest resistances at all immersion times.

The resistance all of the coatings decreases slowly

over an immersion time of 96 hours , which reflected

the stability of the coating [18].

The evolution of a coating resistance is a major

characteristic of the barrier properties of a protective

layer. Generally, the resistance values decreased

during the first hours of immersion, due to the dev-

elopment of conductive pathways inside the film. The

Table 4. Electrical elements fitted values for the samples up to 96 hours of immersion in 0.05 mol L-1 NaCl solution. The error percent-

age associated with each parameter value is given in parenthesis

Time, h Re / cm2 RHF / cm2 CPEHF-Q / µF cm–2 CPEHF-n RCP / cm2 CPECP-Q / nF cm–2 CPECP-n 2103

F1A60M

1 202 (5.7) 92.71 (3.4) 3.29 (5.1) 0.73 (5.8) 134 (5.6) 9.77 (5.7) 0.51 (6.2) 0.79

24 184 (2.7) 69.5 (3.9) 3.69 (4.9) 0.44 (1.4) 89.7 (5.2) 11.4 (2.8) 0.73 (1.5) 1.32

48 167 (2.1) 14.53 (3.5) 4.01 (1.3) 0.41 (3.4) 69.7 (1.2) 13.3 (1.2) 0.69 (0.6) 0.87

96 132 (3.1) 4.31 (5.1) 4.47 (4.2) 0.39 (1.9) 41.8 (1.4) 14.9 (0.4) 0.64 (0.5) 0.95

F2A60M

1 234 (1.7) 357 (2.3) 3.09 (1.8) 0.76 (0.6) - - - 3.17

24 201 (0.8) 124 (1.9) 3.45 (1.4) 0.72 (0.5) - - - 4.71

48 189 (0.6) 64.7 (1.9) 3.73 (1.7) 0.69 (0.2) - - - 2.98

96 166 (0.8) 31.1 (1.2) 3.99 (1.1) 0.66 (0.3) - - - 3.47

F3A60M

1 293 (4.1) 867 (2.5) 0.11 (2.1) 0.84 (1.4) - - - 2.93

24 271 (1.9) 472 (3.7) 0.51 (1.4) 0.81 (0.6) - - - 3.56

48 243 (1.7) 347 (3.9) 1.31 (1.9) 0.79 (0.3) - - - 3.02

96 229 (2.1) 312 (4.2) 2.73 (2.4) 0.76 (0.7) - - - 3.93

F4A60M

1 241 (3.1) 768 (5.7) 2.81 (3.7) 0.78 (0.6) - - - 1.07

24 211 (1.3) 179 (2.7) 2.93 (2.9) 0.73 (0.3) - - - 1.15

48 197 (0.6) 139 (0.8) 3.12 (0.9) 0.71 (0.4) - - - 0.79

96 171 (1.5) 64.7 (3.3) 3.64 (2.3) 0.70 (0.3) - - - 0.69

F5A60M

1 207 (2.2) 63.7 (4.9) 3.61 (4.7) 0.69 (5.4) 912 (4.8) 9.33 (3.2) 0.71 (3.9) 5.31

24 159 (3.5) 52.37 (4.4) 3.94 (5.8) 0.64 (3.8) 287 (3.1) 11.26 (2.7) 0.76 (3.9) 4.39

48 137 (3.1) 24.74 (5.2) 4.42 (5.4) 0.60 (6.6) 146 (4.4) 12.88 (4.1) 0.71 (4.1) 3.27

96 112 (2.9) 8.51 (3.5) 4.89 (2.6) 0.54 (2.2) 61.9 (2.7) 13.86 (2.1) 0.63 (3.4) 2.94


93

high resistance of the sample with 2% plasticizers

should be due to the increasing the ability of polymer

chains to slide over one another; that is to say, the

chain movement enhances the flexibility of the org-

anic radicals that are bound to the silicone atoms.

Figure 6 shows images for all of the hybrid films

after 96 h of the electrochemical impedance tests.

Hybrid films with higher plasticizer concentrations

(F3A60M, Figure 6c and F4A60M, Figure 6d) were

observed to have developed less corrosion products

after the electrochemical tests, confirming the impe-

dance results.

Red corrosion products were observed on the

electrode surface at the end of the experiments, indi-

cating the formation of iron oxides. For hybrid films

without plasticizer (Figure 6e) and with lower plasti-

cizer concentrations F1A60M (Figure 6a) and F2A60M

(Figure 6b), abundant and more localized red cor-

rosion products were observed on their surfaces,

which were considered to be iron oxides.

CONCLUSION

The obtained results are sufficient to state that

all hybrid film coatings shifted potentials indicating

less activity, and diminished the corrosion current

densities (icorr) in relation to the uncoated tinplate (Fl),

evidencing the protection provided by the films. The

hybrid film with 2% of diisodecyl adipate showed the

best performance in the electrochemical impedance

spectroscopy tests.

Films with concentrations lower than 2% plasti-

cizer were thicker, brittle and permeable, which dec-

reases the effect barrier. Films with concentrations

greater than 2% plasticizer featured lower thickness,

irregularities and cracks in the surface, which dec-

reases its efficiency as a barrier film.

Therefore, the addition of 2% of diisodecyl adi-

pate on the silane film improved the morphological

and electrochemical properties of the film, when com-

pared to the film without plasticizer. These results

indicate that this film is a viable alternative to protect

the tinplate against corrosion.

Acknowledgments

This work was done with the financial support of

CAPES, a Brazilian government agency for the

human resources development. Authors would also

like to acknowledge the financial support of CNPq

and FAPERGS, and the Centre for Electron Micro-

scopy of the Federal University of Rio Grande do Sul

for the SEM analysis.

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95

SANDRA RAQUEL KUNST1

LILIAN VANESSA ROSSA

BELTRAMI2

MARIELEN LONGHY1

HENRIQUE RIBEIRO PIAGGIO

CARDOSO2

TIAGO LEMOS MENEZES2

CÉLIA de FRAGA MALFATTI2

1Programa de Pós Graduação em

Engenharia de Processos e

Tecnologias (PGEPROTEC),

University of Caxias do Sul (UCS),

Caxias do Sul, RS – Brasil 2Laboratório de Pesquisa em Corrosão

(LAPEC), Federal University of Rio

Grande do Sul (UFRGS), Porto Alegre,

RS - Brasil

NAUČNI RAD

UTICAJ KONCENTRACIJE DIIZODECIL-ADIPATA U HIBRIDNIM FILMOVIMA NANETIH NA LIMU

Hibridni filmovi na bazi siloksana su krhki i male mehaničke čvrstoce zbog njihovih osobina

staklastog materijala. Stoga je razvijena jedna nova formulacija koja uključuje plastifikator

u cilju povecanja debljine nanosa uniformnog i homogenog hibridnog filma na limu, i

obezbeđenja fleksibilnosti polimernog matriksa. Limene ploče su obložene hibridnim

filmom dobijenim sol-gel postupkom, uz dodatak alkoksidnih prekursora 3-(trimetoksisilil)

propil metakrilat i tetraetokisilan, kao i 0,01 mol/dm3 cerijum nitrata. Ocenjivan je uticaj

koncentracije plastifikatora di-isodecil adipata. Filmovi su okarakterisani skenirajućim

elektronskim mikroskopom, profilometrijom, merenjem potencijala otvorenog kola,

polarizacionim merenjima i spektroskopijom elektrokemijske impedancije. Rezultati su

pokazali da su svi filmovi sa di-isodecil adipatom imali vecu elektrokemijsku performansu u

odnosu na nepremazan lim. Međutim, film sa koncentracijom plastifikatora od 2 % je imao

najbolju performansu u elektrohemijskim testovima, iako je imao manju debljinu.

Ključne reči: hibridni film, diizodecil-adipat, beli lim, elektrohemijsko ponašanje.






97

MILOVAN JANKOVIĆ

SNEŽANA

SINADINOVIĆ-FIŠER

OLGA GOVEDARICA

JELENA PAVLIČEVIĆ

JAROSLAVA

BUDINSKI-SIMENDIĆ

University of Novi Sad, Faculty of

Technology, Novi Sad, Serbia

SCIENTIFIC PAPER

UDC 665.335.2:66.094.3:544.4

https://doi.org/10.2298/CICEQ150702014J

KINETICS OF SOYBEAN OIL EPOXIDATION WITH PERACETIC ACID FORMED IN SITU IN THE PRESENCE OF AN ION EXCHANGE RESIN: PSEUDO-HOMOGENEOUS MODEL

Article Highlights

• In situ epoxidation of soybean oil in the presence of an ion exchange resin is studied

• Occurrence of the reactions during the incremental addition of the reactant is considered

• Pseudo-homogeneous kinetic model is applied

• Temperature dependency of kinetic parameters was determined

• The kinetic model fits well the experimental data

Abstract

A kinetic model was proposed for the epoxidation of vegetable oils with peracetic

acid formed in situ from acetic acid and hydrogen peroxide in the presence of an

acidic ion exchange resin as a catalyst. The model is pseudo-homogeneous with

respect to the catalyst. Besides the main reactions of peracetic acid and epoxy

ring formation, the model takes into account the side reaction of epoxy ring

opening with acetic acid. The partitioning of acetic acid and peracetic acid

between the aqueous and organic phases and the change in the phases’ vol-

umes during the process were considered. The temperature dependency of the

apparent reaction rate coefficients is described by a reparameterized Arrhenius

equation. The constants in the proposed model were estimated by fitting the

experimental data obtained for the epoxidations of soybean oil conducted under

defined reaction conditions. The highest epoxy yield of 87.73% was obtained at

338 K when the mole ratio of oil unsaturation:acetic acid:hydrogen peroxide was

1:0.5:1.35 and when the amount of the catalyst Amberlite IR-120H was 4.04

wt.% of oil. Compared to the other reported pseudo-homogeneous models, the

model proposed in this study better correlates the change of double bond and

epoxy group contents during the epoxidation process.

Keywords: soybean oil, epoxidation, peracetic acid, ion exchange resin, kinetics.

The epoxidation of soybean oil is commercially

important since the obtained epoxide is used as poly-

mer stabilizer and plasticizer, paint and coating com-

ponent, and lubricant. It is also an intermediate for the

production of glycols, alkanolamines, polyols and

polymers [1,2]. Besides performic acid, peracetic acid

is a common oxidizing agent for the epoxidation of

vegetable oils [3,4]. Because of instability and safety

Correspondence: S. Sinadinović-Fišer, University of Novi Sad,

Faculty of Technology, Bul. cara Lazara 1, 21000 Novi Sad,

Serbia. E-mail: [email protected] Paper received: 2 July, 2015 Paper revised: 24 February, 2016 Paper accepted: 11 March, 2016

issues, peracetic acid is usually produced in situ

through the acid-catalyzed peroxidation of acetic acid

with hydrogen peroxide in an aqueous solution [5].

The epoxidation of vegetable oil double bonds with in

situ generated peracetic acid also involves an un-

catalyzed reaction of epoxy ring formation in the org-

anic phase, and a few side reactions of acid-cata-

lyzed epoxy ring opening. Some soluble mineral

acids, like sulphuric acid [1,5-8], and acidic ion

exchange resins can be applied as catalysts for this

process [3,9-20]. The stability of epoxy ring and, thus,

the selectivity of the process, are higher when an ion

exchange resin is used as the catalyst compared to

the mineral acid [21]. Among the sulphonated ion

M. JANKOVIĆ et al.: KINETICS OF SOYBEAN OIL EPOXIDATION… Chem. Ind. Chem. Eng. Q. 23 (1) 97111 (2017)

98

exchange resins investigated, the highest conversion

and selectivity, which were almost constant for five

reuses, were obtained with Amberlite IR-120H [20].

The volume of this resin increases about 43% in

water and about 8% in glacial acetic acid [22]. The

swelling occurs because of the absorption and ads-

orption of solvent molecules on the sulphonated

groups inside the catalyst pores [23,24].

Since the epoxidation of vegetable oils is sig-

nificantly influenced by the reaction conditions, such

as molar ratio of reactants, type of the catalyst, cat-

alyst concentration, temperature, stirring speed, and

presence of the solvent, it is necessary to establish

an appropriate kinetic model for its optimization. In

some studies, the effect of reaction conditions on the

in situ epoxidation of vegetable oils has been inves-

tigated [9,10,18,19], whereas in other studies, the pro-

cess kinetics was also considered [1,3-6,11-17,25-33].

When the ion exchange resin is used as the cat-

alyst for in situ generation of peracetic acid, the react-

ion system for the epoxidation of vegetable oils is

three-phase liquid-liquid-solid (organic-aqueous-solid).

For rigorous mathematical modeling of such a sys-

tem, the intrinsic kinetics, the mass diffusion, and the

partitioning of the components between the phases

have to be considered. Up to the present, this react-

ion system has been described with the pseudo-

-homogeneous [3,16,17] or pseudo-two-phase (liquid-

–solid) [11-15,33] models. Both types of models took

into consideration the main reactions of peracetic acid

and epoxy compound formation, as well as some of

the side reactions of epoxy ring opening that may

occur during the process. The proposed pseudo-

homogeneous models were developed by assuming

that the ion exchange resin is dissolved in the react-

ion mixture. The pseudo-two-phase models were

established by applying Eley-Rideal and Langmuir-

–Hinshelwood-Hougen-Watson approaches to the

reaction of peracetic acid formation. Although a few

authors investigated the partitioning of acetic acid

between the aqueous and organic phases separately

from the epoxidation process [34-36], this pheno-

menon was not considered in the reported kinetic

models.

Because of the aforementioned, the objective of

this study was to develop a kinetic model for the

epoxidation of vegetable oils with peracetic acid

formed in situ in the presence of an acidic ion

exchange resin which takes into consideration the

partitioning of acetic acid and peracetic acid between

the organic and aqueous phases, as well as the

change in the volume of these two phases. The

pseudo-homogeneity of the catalyst with respect to

the aqueous phase is assumed. Besides the main

reactions, the model describes the epoxy ring open-

ing reaction with acetic acid. The model parameters

were determined by fitting the experimental data

obtained for the epoxidation of soybean oil. The pro-

posed model was compared with the pseudo-homo-

geneous model reported in the literature.

EXPERIMENTAL

Materials

Soybean oil was kindly provided by Dijamant

(Zrenjanin, R. Serbia). The acid form of sulphonated

polystyrene-type ion exchange resin Amberlite IR-

120H from Rohm&Hass Co. (Philadelphia, PA, USA)

was used as the catalyst. Glacial acetic acid, 30%

aqueous hydrogen peroxide solution and hydrobromic

acid were purchased from J.T. Baker (Deventer,

Netherlands). Alfapanon (Novi Sad, Serbia) was a

supplier of the aqueous solutions of sodium hydroxide

(0.1 N) and sodium thiosulfate (0.1 N). Iodine (p.a.)

and bromine (p.a.) were purchased from Centrohem

(Stara Pazova, R. Serbia). Potassium iodide (extra

pure), chloroform (min 98.5%), benzene (min 99.8%),

potassium hydrogen phthalate (min 99.0%) were

bought from LachNer (Neratovice, Czech Republic).

Hydrogen bromide solution (33.0%) in acetic acid and

crystal violet were purchased from Sigma-Aldrich (St.

Louis, MO, USA).

Epoxidation procedure

The epoxidation of soybean oil in bulk was car-

ried out with peracetic acid formed in situ according to

the method reported in the literature [15]. Soybean oil

with an initial iodine number (IN0) of 128.62, which

corresponds to 0.5067 mol of double bond per 100 g

of oil, was used for the syntheses. The molar ratio of

soybean oil unsaturation:acetic acid:hydrogen per-

oxide was approximately 1:0.5:1.35 for all runs. The

oil (approximately 150 g) was mixed with an appro-

priate mass of glacial acetic acid before being poured

into a 1000 mL three-neck glass reactor equipped

with a magnetic stirrer, a thermometer, a reflux con-

denser and an addition funnel. The reactor was

placed in a water bath. Amberlite IR-120H ion

exchange resin was introduced to the reactor before

the reactants. The amount of catalyst was 1.98–

–7.86 wt.% of oil. Subsequently, the 30% aqueous

hydrogen peroxide solution was added drop-wise to

the reaction mixture at a constant rate within 40 to 65

min. During the addition, the temperature of the mix-

ture was maintained at 323 or 338 K with fluctuation

of less than ±1 K. Further, where applicable, the tem-


99

perature was increased at a uniform rate to the

desired level and maintained at 323, 338, 348 or 353

K within ±1 K. The temperature of the reaction mix-

ture was controlled by changing the water bath tem-

perature. The fine dispersion of oil in the reaction

mixture was achieved with the uniform agitation using

a PTFE-coated cylindrical stirring bar (35 mm6 mm)

under constant stirring speeds of 900, 1000 or 1100

rpm. The beginning of the addition of hydrogen per-

oxide solution was considered to be the “zero time” of

the process. The progress of epoxidation was fol-

lowed by withdrawing 8 mL samples of the reaction

mixture at defined time intervals. The first sample was

taken immediately after the completion of the hydro-

gen peroxide addition, whereas the second imme-

diately after raising the reaction temperature, where

applicable. The others were collected at about 30, 60

or 120 min intervals. After cooling and centrifugation

of the sample, the separated organic phase was

washed with water (323 K) until pH 7. Water was eva-

porated from the sample at 333 K under the vacuum.

The evaporation lasted a minimum of 1 h. The

samples were then analyzed to determine the iodine

number (IN) and epoxy oxygen content (EO). The

infrared analysis of the samples was also provided.

Analyses

The iodine number and epoxy oxygen content

were measured in triplicate according to the Hanus

method and the standard HBr-acetic acid method,

respectively [37]. Using attenuated total reflectance

(ATR) FT-IR spectroscopy, the spectra were recorded

by Thermo Finnigan’s Nicolet 5700 FTIR spectro-

meter in the range of 4000-400 cm-1 by accumulating

32 scans at a resolution of 4 cm-1.


Soybean oil epoxidation

Nine epoxidation runs were conducted using

soybean oil and peracetic acid formed in situ from

acetic acid and 30% hydrogen peroxide aqueous

solution in the presence of Amberlite IR-120H. The

mole ratio of soybean oil unsaturation:acetic acid:hyd-

rogen peroxide was approximately 1:0.5:1.35 for all

runs. The reaction conditions and some of the results

are summarized in Table 1. Residual iodine number

(IN), conversion of double bond (X), relative epoxy

yield (REY) and selectivity (SE) are presented only for

the content of epoxy oxygen (EOt) reached in each

run after defined period of time (t).

The disappearance of double bonds and form-

ation of epoxy groups during the runs were monitored

on the basis of the FT-IR spectrum of the samples, as

illustrated in Figure 1. The changes in the intensities

of epoxy group doublet band (with maxima at 823 and

845 cm–1) and double bond band (at 3007 cm-1) were

qualitatively analyzed.

The progress of the epoxidation was quantified

by determining the iodine number (IN) and epoxy

oxygen content (EO) for all samples withdrawn during

the run. The values are presented as points in Figures

2-4.

There are detailed discussions in our previous

studies [15,16] and in other studies [3,9–14,17] on the

Table 1. Reaction conditions and the residual iodine number (IN), conversion of double bond (X), relative epoxy yield (REY) and sel-

ectivity (SE) for content of epoxy oxygen (EOt) reached after reaction time (t) in each run of the epoxidation of soybean oil (SO)(a) in bulk

with peracetic acid formed in situ from 30% hydrogen peroxide aqueous solution (aqHP) and acetic acid (A) in the presence of Amberlite

IR-120H as the catalyst when the molar ratio of soybean oil unsaturation:acetic acid:hydrogen peroxide was approximately 1:0.5:1.35; initial iodine number IN0 = 128.62 corresponds to theoretical epoxy oxygen content (EOth) of 7.50% where EOth = 100{(IN0/2AI)/

/[100+(IN0/2AI)AO]}AO

Run Measured mass, g Amberlitea

wt.%

Stirring speed

rpm

Temperature, K t

min

EOt

% INb

Xc)

%

REYd

% SEe

SO A aqHP H2O2 addition Reaction

1 150.00 22.82 116.83 3.92 [4.21] 1000 323 323 630 4.65 44.2 65.64 62.00 0.94

2 150.00 22.82 116.83 3.92 [4.21] 1000 323 338 645 6.25 7.45 94.21 83.33 0.88

3 150.00 22.82 116.83 3.92 [4.21] 1000 323 348 635 6.22 1.73 98.65 82.93 0.84

4 150.00 22.82 116.83 3.92 [4.21] 1000 323 353 470 6.08 4.45 96.54 81.06 0.84

5 145.44 21.66 115.90 4.04 [4.27] 1100 338 338 625 6.58 1.99 98.45 87.73 0.89

6 150.23 22.81 109.46 3.91 [4.45] 1000 338 338 480 6.45 7.22 94.39 86.00 0.91

7 145.50 21.91 117.02 4.04 [4.23] 900 338 338 615 6.40 2.71 97.89 85.33 0.87

8 148.32 21.84 112.56 1.98 [2.19] 1100 323 353 590 6.37 2.87 97.77 84.93 0.87

9 149.82 21.57 114.53 7.86 [8.65] 1100 323 323 630 5.51 22.63 82.41 73.46 0.89

aCatalyst concentration is expressed in percentage of soybean oil weight; value given in square brackets is percentage of catalyst in respect to acetic acid

and 30% hydrogen peroxide weight; bresidual iodine number, IN; cdouble bond conversion, X = 100(IN0-IN)/IN0; drelative epoxy yield, REY = 100EOt/EOth; eselectivity, SE = EOtIN0/[EOth(IN0-IN)]


100

Figure 1. FT-IR spectra of soybean oil and samples taken after 55, 175 and 625 min of soybean oil epoxidation in bulk with peracetic

acid formed in situ from acetic acid and hydrogen peroxide in the presence of amberlite IR-120H in the amount of 4.04 wt.% of oil at 338

K and 1100 rpm, when the mole ratio of double bond in oil:acetic acid:hydrogen peroxide was approximately 1:0.5:1.35.

Figure 2. Time dependency of experimentally determined (points) and calculated (curves) iodine number (IN) and epoxy oxygen content

(EO) for the epoxidation of soybean oil in bulk with peracetic acid generated in situ in the presence of Amberlite IR-120H in the amount

of approximately 4.0 wt.% of oil at 338 K, when the stirring speeds (S) were 900 and 1100 rpm and when the mole ratio of double bond

in oil:acetic acid:30 wt.% hydrogen peroxide was approximately 1:0.5:1.35.

Figure 3. Time dependency of experimentally determined (points) and calculated (curves) iodine number (IN) and epoxy oxygen content

(EO) for the epoxidation of soybean oil in bulk with peracetic acid generated in situ in the presence of Amberlite IR-120H in the amount

of 1.98, 3.92 and 7.86 wt.% of oil at 323 (Runs 1and 9) or 353 K (Run 8), when the mole ratio of double bond in oil:acetic acid:30 wt.%

hydrogen peroxide was approximately 1:0.5:1.35.


101

Figure 4. Time dependency of experimentally determined (points) and calculated (curves): a) iodine number (IN) and b) epoxy oxygen

content (EO), for the epoxidation of soybean oil in bulk with peracetic acid generated in situ in the presence of Amberlite IR-120H in the

amount of 3.92 wt.% of oil at 323, 338, 348 and 353 K and 1000 rpm, when the mole ratio of double bond in oil:acetic acid:30 wt.%

hydrogen peroxide was approximately 1:0.5:1.35.

influence of the reaction conditions on the conversion

of double bond, relative epoxy yield, and selectivity

for the in situ epoxidation of vegetable oils with pera-

cetic acid in the presence of the ion exchange resin.

This study discusses the results obtained for the in

situ epoxidation of soybean oil.

The stoichiometric ratio of hydrogen peroxide to

vegetable oil unsaturation is 1:1. However, this oxy-

gen agent is usually applied in excess for the epoxid-

ation of vegetable oils [3,9-18]. The mole ratio of

hydrogen peroxide to double bond was studied in the

range from 0.8 to 3 for various oils [3,9-18]. The

significant increase in the rate of epoxy ring formation

was observed with an increase in molar ratio from 0.8

to 1.5 [9,12-14] or 2 [10,11]. With a further increase in

molar ratio, there was no appreciable decrease of the

residual iodine number or change of the epoxy oxy-

gen content [9-14]. Also, when a mole ratio higher

than 1.5 was applied, a higher rate of the epoxy ring

opening reactions was observed [3,9,10]. Addition-

ally, the excess of hydrogen peroxide solution in the

reaction system decreases the concentration of acetic

acid as the other reactant in the reaction of peracetic

acid formation. Hence, a hydrogen peroxide to veget-

able oil unsaturation molar ratio of about 1.35 was

chosen for this study.

The stirring speed influences the external mass

transfer in the multiphase reaction systems. To red-

uce the mass transfer resistance, adequate mixing

has to be achieved. Under a particular stirring speed,

the intensity of mixing also depends upon the type

and design of a stirrer [18]. At higher stirring speeds,

the mass transfer resistance is lower. Consequently,

the rate of the epoxidation should be higher. How-

ever, no significant influence on the double bond con-

version and relative epoxy yield was observed when

the stirring speed changed from 900 to 1100 rpm

under the conditions applied in this study (Figure 2).

The reported amounts of catalyst for the epoxid-

ation of vegetable oils vary from 1.28 to 25 wt.% of oil

[3,9-18]. The quantity of Amberlite IR-120H used in

this study was 1.98–7.86 wt.% of oil. To ensure good

mixing of the reaction mixture, such low amounts of

the catalyst were chosen. As concluded in most

papers, when a higher amount of the catalyst was

used, an increase in the double bond conversion and

epoxy yield was observed. At the same time, the

selectivity of the process was reduced with an inc-

rease in the catalyst load (Table 1, runs 1 and 9). This

was the consequence of the promotion of acid-cat-

alyzed epoxy ring opening reactions. Figure 3 shows

the change in iodine number and epoxy oxygen con-

tent over time for the epoxidations of soybean oil per-

formed under different catalyst concentrations. It

should be noted that the reaction temperature of the

run with the lowest catalyst loading was higher than

the temperature for the other two runs.

The temperature of the process should be

adjusted to increase the rate of the peracetic acid

formation and epoxidation reaction, but not to com-

promise the stability of the epoxy ring. For the epoxid-

ations of different vegetable oils, the temperatures

varied from 303 to 363 K [3,9-17]. In order to deter-

mine the dependency of the Arrhenius type model

parameters from temperature, the epoxidations of

soybean oil were carried out at 323, 338, 348 and 353

K in this study. The reaction temperatures were the

same or higher than the temperature at which the


102

hydrogen peroxide solution was added into the react-

ion mixture. The addition was performed gradually to

avoid an increase in temperature due to the exo-

thermic effect of the reaction of peracetic acid form-

ation. An increase in the reaction rates with an inc-

rease in the reaction temperature was observed

(Figure 4a and b). Consequently, the reaction time for

achieving the same epoxy yield was shortened.

The highest yield of epoxide of 87.73% with

almost complete conversion of double bonds of

98.45% was achieved after 625 min, when the tem-

perature of the hydrogen peroxide solution addition

was the same as desired reaction temperature of 338

K. The amount of the catalyst was only 4.04 wt.% of

oil. The selectivity of double bond conversion to

epoxide of 0.89 and the residual unsaturation of epox-

idized soybean oil of 1.99 imply good quality of the

obtained product (Table 1).

Kinetic model

The epoxidation of soybean oil involves an acid-

catalyzed reaction of the peracetic acid (P) and water

(W) formation from acetic acid (A) and hydrogen

peroxide (H):

(1)

followed by an un-catalyzed reaction of the double

bond (D) conversion into the epoxy ring (E):

(2)

Among a few acid-catalyzed side reactions of

epoxy ring cleavage (Eqs. (3a)-(3e)), the most likely is

the reaction with acetic acid that leads to the form-

ation of hydroxy acetate (HA), Eq. (3a) [38]:

(3)

In this three-phase organic-aqueous-solid react-

ion system, the catalyst Amberlite IR-120H swells due

to its sorption affinity towards the polar components. It

can be assumed that the swelling of the resin is com-

pleted already at the beginning and that the swelling

degree is constant during the epoxidation. Because of

the preferential sorption of some components over

the others, the concentrations of the components dif-

fer in the aqueous bulk phase and inside the catalyst

pores [23,24]. The proceeding of the reactions causes

the mass transfer phenomena across the organic-

aqueous interface, across the aqueous-solid inter-

face, and through the resin pores. In Figure 5a, the

system phases with the possible reactions and mass

transfer of the components across the reaction sys-

tem are schematically presented.

Acetic acid and hydrogen peroxide diffuse from

the aqueous bulk phase to the catalyst external sur-

face and further into the catalyst pores, in both cases,

to reach the active sites where they react. The pro-

ducts of the reaction, peracetic acid and water, mig-

rate to the aqueous phase. As a consequence of the

reaction, a concentration gradient of the components

exists inside the catalyst pores. From the aqueous

phase, the peracetic acid diffuses into the organic

bulk phase where it epoxidizes the triglyceride double

bond. In the epoxidation reaction, regenerated acetic

acid diffuses to the aqueous phase. A fraction of the

acid reacts with the epoxy ring, however. The epoxid-

ized triglycerides diffuse from the organic phase to

the aqueous phase to reach the catalyst. Due to steric

hindrance, the epoxidized triglycerides cannot enter

the catalyst pores. However, they react with acetic

acid or hydrogen peroxide at acidic active sites avail-

able at the external surface of the catalyst [39]. Tri-

glycerides with opened epoxy rings diffuse into the

aqueous and further to the organic phase. The epoxy

ring opening reaction also occurs at the organic-aque-

ous interface according to some authors [25,29].

Since a mathematical description of the three-

-phase reaction system for the epoxidation of veget-

able oils is complex, the kinetic models developed in

other studies are based on a simplified reaction

scheme, a reduced number of phases, and/or neg-

lected transport phenomena [11-17,26,33]. The model

proposed in this study is established by assuming that

only reactions (1), (2) and (3a) occur during the epox-

idation. Some authors reported that the effect of

internal mass transfer is negligible in the reaction sys-

tem for the epoxidation of vegetable oils with pera-

cetic acid formed in situ [11-14]. The same was con-

cluded for the peracetic acid formation investigated

apart from the epoxidation process [30]. Therefore,

the preferential sorption and internal mass transfer

inside the resin pores were neglected in this study.

This enabled using a pseudo-homogeneous model to

describe the kinetics of the reaction of peracetic acid

formation.


103

To develop the mathematical model for the

epoxidation of vegetable oils, the material balances

for components in the reaction system were derived

from the equation for semi-batch, i.e., feed-batch

reactor [40]:

,

1

d

d

NRj

j j i i

i

NF V r

t (4)

where Nj (mol) is the number of moles of component j;

t (min) indicates the reaction time; Fj (mol/min) is the

molar flow of component j; V (L) indicates the volume;

NR is the total number of reactions; j,i is the stoi-

chiometric coefficient of component j in the reaction i;

and ri is the rate of reaction i.

It is known from the literature that the solubility

of hydrogen peroxide in the organic phase of the

epoxidation reaction system can be neglected [39].

The amount of water in the organic phase of the

epoxidized soybean oil-acetic acid-water system

ranges from 1.62 to 3.04 wt.% at 323-353 K [36],

whereas it is not present in the organic phase of the

soybean oil-acetic acid-water system [34,35]. How-

ever, for the sake of simplicity, it was assumed that

the water is present only in the aqueous phase,

whereas the triglycerides are present only in the org-

anic phase of the investigated reaction system. Thus,

the number of moles of the components is: aqH HN N ,

aqW WN N , o

D DN N , oE EN N and o

HA HAN N . Since

acetic acid and peracetic acid are partitioned between

the organic and aqueous phases, the mass transfer

and concentrations of these components in the par-

ticular phase should be taken into consideration [5]:

aq

aqaq aq aq o aq oA

A, L,A A A A

1

d

d

NR

i i

i

NV r k a C K C V

t (5)

o

oo o o o aq oA

A, L,A A A A

1

d

d

NR

i i

i

NV r k a C K C V

t (6)

aq

aqaq aq aq aq o oP

P, L,P P P P

1

d

d

NR

i i

i

NV r k a K C C V

t (7)

o

oo o o aq o oP

P, L,P P P P

1

d

d

NR

i i

i

NV r k a K C C V

t (8)

where superscripts aq and o denote the aqueous and

organic phases, respectively; kL,A and kL,P (m/min) are

Figure 5. The reaction system for the epoxidation of soybean oil with peracetic acid formed in situ from acetic acid and hydrogen

peroxide in the presence of Amberlite IR-120H: a) reactions and mass diffusion and b) in this work considered reactions

and partitioning of the components.


104

the mass transfer coefficients for acetic acid and

peracetic acid, respectively; a (m2/m3) indicates the

interface area; oAC , aq

AC and oPC , aq

PC (mol/L) are the

concentrations of acetic acid and peracetic acid in a

particular phase, respectively; and KA and KP indicate

the partition coefficients for acetic acid and peracetic

acid, respectively.

The partition coefficients of acetic acid and per-

acetic acid are defined as follows:

o o oA A

A aq aq aqA A

/

/

C N VK

C N V (9)

o o oP P

P aq aq aqP P

/

/

C N VK

C N V (10)

When the number of moles of acetic acid in the

organic phase is expressed from Eq. (9) and substi-

tuted in the material balance equation for acetic acid:

aq oA A AN N N (11)

the following equation is derived:

aq o

aq A AA A aq

K N VN N

V (12)

From Eq. (12), the number of moles of acetic

acid in the aqueous phase may be expressed as a

function of the total number of moles of acetic acid in

the system:

aqaq AA aq o

A

V NN

V K V (13)

Therefore, the concentration of acetic acid in the

aqueous phase is:

aq AA aq o

A

NC

V K V (14)

Likewise, the number of moles and concentra-

tion of acetic acid in the organic phase are as follows:

oo A AA aq o

A

K V NN

V K V (15)

o A AA aq o

A

K NC

V K V (16)

Analogously, the concentrations of peracetic

acid in the aqueous and organic phases can be

derived and expressed as follows:

aq PP aq o

P

NC

V K V (17)

o P PP aq o

P

K NC

V K V (18)

To reduce the external mass transfer resistance

between the organic and aqueous phases, as well as

between the aqueous and solid phases, effective mix-

ing was ensured by vigorous stirring in this study. By

investigating the influence of the stirring speed on the

kinetics of the epoxidation process, it was found that

the mass transfer is faster than the reaction kinetics.

Thus, the mass transfer resistance can be neglected

and the terms related to the mass transfer can be

omitted from the kinetic model [4]. The material bal-

ance equation for acetic acid, on the basis of the Eqs.

(5), (6) and (11), becomes:

aq o

aq oA A A

aq aq aq o o oA, A,

1 1

NR NR

i i i i

i i

dN dN dN

dt dt dt

V r V r

(19)

and, likewise, the material balance equation for per-

acetic acid is as follows:

aq o

aq oP P P

aq aq aq o o oP, P,

1 1

NR NR

i i i i

i i

dN dN dN

dt dt dt

V r V r

(20)

Now, the mathematical model that describes the

reaction system for the epoxidation of vegetable oils

is derived as:

3

aq aq aqaq aq aq aqH P W

1 A HH O11

d

d

N C Ck C C C V

t K (21)

aq aqH H

H

1

d d

d d

N NF

t t (22)

aqo o o o o oA H

2 P D 3 E A

1

d d

d d

nN Nk C C V k C C V

t t (23)

aqo o oP H

2 P D

1

d d

d d

N Nk C C V

t t (24)

aq aqW H

W

1

d d

d d

N NF

t t (25)

o

o o oD2 P D

d

d

Nk C C V

t (26)

o

o o o o o oE2 P D 3 E A

d

d

nNk C C V k C C V

t (27)


105

o

o o oHA3 E A

d

d

nNk C C V

t (28)

where k1 (mol-2min-1), k2 (mol-1min-1) and k3 (mol-n min-1)

are the rate coefficients for the reactions (1), (2) and

(3a); and n is the order of epoxy ring opening reaction

with respect to acetic acid which was assumed to be

the first or the second. The reactions and the parti-

tioning of the components considered in the model

proposed in this study are shown in Figure 5b.

Since the model is pseudo-homogeneous with

respect to the aqueous phase, the concentration of

hydronium ions ( +3H O

C ) corresponds to the complete

dissociation of resin sulphonated groups in the aque-

ous phase and it is expressed as:

3

aq saqH O

mCC

V (29)

where m (g) is the mass of the catalyst; and Cs (mol/g

catalyst) indicates the concentration of active catalyst

sites (sulphonated groups).

After substituting the expressions (29), (14) and

(16)-(18) in the model Eqs. (21)–(28) and expressing

the concentrations of components as the ratio of the

number of moles and the phase volume, the model

equations become functions of the total number of

moles of the components:

H

1

s 1 A H P Waq aq o aq o

A 1 P

d

d

( )

N

t

mC k N N N N

V V K V K V K V

(30)

H HH

1

d d

d d

N NF

t t (31)

A H

1

P P D A A2 3 Eaq o aq o

P A

d d

d d

n

N N

t t

K N N K Nk k N

V K V V K V

(32)

P H P P D2 aq o

P1

d d

d d

N N K N Nk

t t V K V (33)

W HW

1

d d

d d

N NF

t t (34)

D P P D2 aq o

P

d

d

N K N Nk

t V K V (35)

E P P D A A2 3 Eaq o aq o

P A

d

d

nN K N N K N

k k Nt V K V V K V

(36)

HA A A3 E aq o

A

d

d

nN K N

k Nt V K V

(37)

Instead of using the number of moles of com-

ponent j (Nj) in the model Eqs. (30)-(37), the amount

of component j in the reaction system can be expres-

sed as the number of moles of component j per 100 g

of oil ([j]).

For the regression of the experimental data by

the proposed model, some approximations were

applied as follows.

The volumes of the organic and aqueous phases

were accepted as the sums of the components’

volumes. For this purpose, the influence of tempe-

rature on the densities of both phases was neglected.

The volume of the aqueous phase may be expressed

per 100 g of oil as:

aq aq aq aq aqW H A Pv v v v v (38)

The volume of each component in the aqueous

phase per 100 g of oil was expressed as the ratio of

mass and density of the component. Further, the

mass was defined via the component’s molecular

mass and its amount, or its concentration in the

aqueous phase:

aq aq aq aqaq W H A A P P

W H A P

[W]M [H]M M MC v C vv (39)

where W , H , A and P (g/L) are densities of

water, hydrogen peroxide, acetic acid and peracetic

acid, respectively. By substituting the concentrations

of acetic acid and peracetic acid, expressed from

Eqs. (14) and (17), into Eq. (39), the volume of the

aqueous phase becomes:

The volume of the organic phase per 100 g of oil

is calculated from the material balance of the initial

double bonds, which undergo epoxidation and further

the acetylation, as follows:

aq aqaq W H A P

o aq o aqW H A A P P

[W]M [H]M [A] M [P] M

( ) ( )

v vv

K v v K v v (40)

o o o oo 0 0A A P P

oA P o 0

100[D] (100 76.054[D] )[HA] (100 16[D] )[E]M M

[D]

C v C vv (41)


106

where [D]0 (mol/100 g of oil) is the initial amount of

double bond in the oil; and oo (g/L) is the density of

the oil in the organic phase. Further, density of the oil

in the organic phase was calculated by the linear

interpolation of densities of soybean oil ( SO ) and

epoxidized soybean oil ( ESO ) for the particular double

bond and epoxy group concentrations in the oil:

oo SO ESO SO

ESO

EO

EO (42)

where EOESO (wt.%) is the epoxy oxygen content in

epoxidized soybean oil. By substituting the concentra-

tions of acetic acid and peracetic acid in the organic

phase, expressed on the basis of Eqs. (16) and (18),

into the Eq. (41), the volume of the organic phase

becomes:

The system of two nonlinear Eqs. (40) and (43),

with two unknowns, vaq and vo, may be solved by

some of numerical methods, such as the Newtonian

method of simultaneous solving the system of non-

linear equations. However, in this case, the sequential

determination of the unknowns’ values with a two

loops algorithm was applied to simplify the calcul-

ation. The value of vo was determined in the outer

loop, whereas the value of vaq was determined in the

inner loop for the current value of vo. In both loops,

the modified Newtonian method for nonlinear equa-

tions was applied. The calculation of the phases' vol-

ume must be run at each step of integration of the

differential Eqs. (30)-(37).

The amounts of acetic acid and peracetic acid in

one phase were calculated on the basis of the parti-

tion coefficients and known amounts of components

in the other phase. The partition coefficient for acetic

acid at temperature T was approximated by double

linear interpolation of the values of the partition coef-

ficient for acetic acid for the soybean oil-acetic acid-

–water and epoxidized soybean oil-acetic acid-water

systems:

A,SO, 1 A,SO,

A,SO A,SO,

1

p pp p

p p

K KK K T T

T T (44)

A,ESO,p 1 A,ESO,

A,ESO A,ESO,

1

pp p

p p

K KK K T T

T T (45)

A,SO A,ESO

A

D E

D E

K KK (46)

where KA,SO, KA,SO, p and KA,SO,p+1 are the partition

coefficients for acetic acid between the soybean oil

and water at the temperatures T, Tp and Tp+1, res-

pectively; KA,ESO, KA,ESO,p and KA,ESO,p+1 are the parti-

tion coefficients for acetic acid between the epox-

idized soybean oil and water at the temperatures T,

Tp and Tp+1, respectively; and Tp and Tp+1 are the tem-

peratures given in Table 2 with the closest values to

the T. The values of the partition coefficients for these

three-component systems were calculated on the

basis of the experimental data given in the literature

[34,36]. Further, it was assumed that the value of the

partition coefficient for peracetic acid is always 2.5

times higher than the value of the partition coefficient

for acetic acid [5].

Table 2. Values of the partition coefficient for acetic acid in the

systems soybean oil-acetic acid-water (SO-A-W) and epoxid-

ized soybean oil-acetic acid-water (ESO-A-W) calculated on the

basis of experimental data reported the literature [34,36]

T / K Partition coefficient for acetic acid

SO-A-W ESO-A-W

323 0.04255 0.1647

338 0.04568 0.1620

353 0.04867 0.1860

For the temperature dependency of the chem-

ical equilibrium constant for peracetic acid formation

from acetic acid and hydrogen peroxide in an aque-

ous solution (K1), the expression reported in the

literature was applied [41]:

1

6 2

exp(12.2324ln 0.0229913

9.70452 10 3045.76 / 72.8758)

K T T

T T (47)

where T (K) is the temperature.

A drop-wise addition of the hydrogen peroxide

solution to the reaction mixture is approximated with

continuous flows of hydrogen peroxide (FH) and water

(FW), both in (mol/(min·100 g oil)):

1 1HS H H HS HS

H

HS

0

m w M t t tF

t t (48)

1 1HS H W HS HS

W

HS

(1 )

0

m w M t t tF

t t (49)

o oo 0 0A P

oaq aq

o 0o oA P

PA

100[D] (100 76.054[D] )[HA] (100 16[D] )[E][A] M [P] M

[D]

v vv

V vv v

K K

(43)


107

where mHS (g) is the mass of the hydrogen peroxide

aqueous solution; wH is the mass fraction of hydrogen

peroxide in the solution; MH and MW (g/mol) are mol-

ecular masses of hydrogen peroxide and water; and

tHS (min) is the duration of hydrogen peroxide solution

addition.

The temperature dependency of the rate coef-

ficient ki for the reaction i is expressed with repara-

meterized form of the Arrhenius equation [42]:

a,E

,0

1 1exp

i

i i

a

kk k

R T T (50)

where ki,0 and a,Eik are the modified Arrhenius

equation constants related to the frequency coef-

ficient and the activation energy, respectively; R

(8.3143 J/(mol K)) is the universal gas constant; and

Ta (K) indicates an average temperature of experi-

ments.

Kinetic parameters

Prior to fitting the experimental data with the

proposed model, it is necessary to define the time

dependency of the reaction temperature. In this study,

the hydrogen peroxide aqueous solution addition to

the reaction mixture was isothermal. The increase in

temperature to the desired reaction temperature was

approximated as linear with reaction time. Further, the

epoxidation was run isothermally. These temperature

changes are described mathematically as follows:

where THS (K) is the temperature of the hydrogen

peroxide solution addition; Tr (K) is the temperature of

the reaction; and tTI (min) is the period of the

temperature increase.

The constants of the reparameterized Arrhenius

equation, ki,0 and a,Eik , were determined by fitting the

values of double bond [D] and epoxy group [E]

amount changes with reaction time (t) for the soybean

oil epoxidations run under different reaction condi-

tions. The Marquardt method was used to fit the data

[43]. The following objective function (F) was mini-

mized:

NSNRN 2 2calc exp calc exp

, , , ,1 1

F D D E Ek

k l k l k l k lk l

(52)

where NRN indicates the number of runs and NSk is

the number of samples in run k; [D] and [E] are

determined as [D] = IN/[2AI] and [E] = EO/[100AO],

where AI is the atomic mass of iodine and AO is the

atomic mass of oxygen; and superscripts calc and

exp indicate the calculated and experimentally deter-

mined value, respectively.

The model’s system of differential Eqs. (30)-(37)

was integrated by applying a fourth order Runge-

–Kutta method.

The model parameters, i.e., the rate coefficients

ki for the investigated reactions were calculated when

the average temperature of the experiments (Ta) was

accepted as 338 K and assuming the first and the

second order of the epoxy ring opening reaction with

respect to acetic acid. The results are presented in

Table 3 together with the values of the least sum of

squares (objective function F) and average absolute

deviations for amounts of double bond (AAD[D]) and

epoxy group (AAD[E]).

The least sum of squares is lower (0.06519)

when the order of the side reaction was assumed as

the first. The temperature dependencies of the kinetic

parameters for such assumption are determined as

follows:

1

83959.13 1 1exp 10.15365s

a

k CR T T

(53)

2

1425.418 1 1exp 2.564719

a

kR T T

(54)

3

6873.391 1 1exp 6.143859

a

kR T T

(55)

Note that the rate coefficients are not intrinsic

but apparent, since the effect of the mass transfer

resistance on the kinetics was not taken into con-

sideration. Also, the rate coefficient for peracetic acid

formation is expressed in combination with the con-

centration of the active catalyst sites (Cs). An increase

of all rate coefficients with an increase in temperature

was obtained. Since all calculated reactant and pro-

duct amounts have positive values, the viability of the

developed kinetic model is confirmed. Due to lower

average absolute deviation of 0.01258 than 0.01530,

it can be concluded that the model fits the time vari-

ation of double bond amount slightly better than

epoxy group amount, respectively (Table 3). On the

basis of double bond and epoxy group amounts

HS HS

HS r HS HS TI TI HS, HS TI

HS TIr

( )( ) / ( )

( )

k

T t t

T T T T t t t t t t t t

T t t t

(51)


108

estimated with developed model, the iodine number

and epoxy oxygen content, respectively, were calcul-

ated. In Figures 2-4 are shown changes of the iodine

number and epoxy oxygen content over time for the

epoxidations of soybean oil with peracetic acid gener-

ated in situ in the presence of Amberlite IR-120H. The

proposed model fits the corresponding experimental

data reasonably well.

Comparison of the proposed model with the

pseudo-homogeneous model reported in the literature

The kinetic model proposed in this study is

established assuming the pseudo-homogeneity of the

catalyst with respect to the aqueous phase. Unlike the

other pseudo-homogeneous models found in the

literature [16,31,44], it takes into consideration the

partitioning of acetic acid and peracetic acid between

the organic and aqueous phases, as well as the

changing of the phases’ volumes during the process.

The model also takes into account the occurrence of

the reactions during the incremental addition of the

oxidizing agent solution and defined temperature

changes during the epoxidation process. These phen-

omena were already considered in the pseudo-homo-

geneous model developed for the in situ epoxidation

of castor oil [16]. In order to compare the latter

pseudo-homogeneous model with the model pro-

posed in this study, the experimental data obtained

for the epoxidation of soybean oil under defined react-

ion conditions (Table 1) were fitted. In both models,

the beginning of the addition of hydrogen peroxide

solution to the reaction mixture was considered as the

zero reaction time. Also, the variation of the chemical

equilibrium constant for peracetic acid formation with

temperature was defined with the same expression.

Further, the temperature dependencies of the kinetic

parameters were defined with the same Arrhenius

type model. The same objective function F was applied

for both regressions. The least sum of squares of

0.06519 obtained for the model developed in this

study is more than 44% lower than those of 0.11800

obtained for the reported model when the second

order of the epoxy ring opening reaction with respect

to acetic acid was assumed (Table 3). This confirmed

that the improved model better describes the reaction

system for the epoxidation of soybean oil conducted

under the investigated conditions. Since it was shown

in the literature [16] that the model developed for the

epoxidation of castor oil fits the experimental data

better than the two pseudo-homogeneous kinetic

models proposed by other authors [31,44], it can be

concluded that the model proposed in this study cor-

relates the change of double bond and epoxy group

contents during the in situ epoxidation of vegetable

oils better than the other reported pseudo-homo-

geneous models.

Table 3. Statistical values of the model parameters determination when the models proposed in this study and in the literature [16] were

applied for the in situ epoxidation of soybean oil. The order of the epoxy ring opening reaction with respect to acetic acid was 1 or 2. The

constants of the reparameterized Arrhenius equation for compared models are given

Parameter

This study Reference [16]

Order of the epoxy ring opening reaction, n

1 2 1 2

Error

F 0.06519 0.06764 0.12250 0.11800

AAD[D]a 0.01258 0.01275 0.01999 0.01953

AAD[E]b 0.01530 0.01556 0.01937 0.01879

Constants of the reparameterized Arrhenius equation

(k1Cs)0 -10.15365 -10.20897 -5.171507 -5.172997

(k1Cs)Ea 83959.13 82829.17 79854.73 79039.30

k2,0 -2.564719 -2.501162 -1.332284 -1.282917

k2,Ea 1425.418 -1040.259 -11635.34 -13083.98

k3,0 -6.143859 -4.765239 -5.737784 -4.010654

k3,Ea 6873.391 300.7545 -3704.474 -3174.165

aAverage absolute deviation for double bond amount:

NSNRNcalc exp

[D] NRN , ,1 1

1

1AAD D D

NS

k

k l k lk l

kk

baverage absolute deviation for epoxy group amount:

NSNRNcalc exp

[E] NRN , ,1 1

1

1AAD E E

NS

k

k l k lk l

kk


109

CONCLUSION

The epoxidations of soybean oil with peracetic

acid formed in situ from acetic acid and hydrogen

peroxide in the presence of an acidic ion exchange

resin as the catalyst were carried out under different

temperatures of incremental oxidizing agent addition,

reaction temperatures, catalyst concentrations and

stirring speeds. It was determined that the highest

epoxide yield was achieved at the reaction tempe-

rature of 338 K after 625 min when approximately 0.5

mol of glacial acetic acid and 1.35 mol of 30%

aqueous hydrogen peroxide solution per mole of

double bond of soybean oil were used. The epox-

idation was catalyzed with Amberlite IR-120H in the

amount of only 4.04 wt.% of oil. A kinetic model was

developed assuming the pseudo-homogeneity of the

catalyst with respect to the aqueous phase. The pro-

posed model took the partitioning of acetic acid and

peracetic acid between the organic and aqueous

phases into consideration, as well as the changing of

the phases’ volumes during the process. The tempe-

rature dependency of the kinetic parameters was det-

ermined and an increase in all reaction rate coef-

ficients with an increase in temperature was obtained.

A comparison was made between the proposed

model and the pseudo-homogeneous model reported

in the literature. The model developed in this study fits

the experimental data obtained for the in situ epox-

idations of soybean oil run under the investigated

conditions better than the kinetic model taken for the

comparison.

Acknowledgements

This work is part of the Project No. III45022

financially supported by the Ministry of Education,

Science and Technology Development of the Rep-

ublic of Serbia. The authors thank the German Aca-

demic Exchange Service (DAAD) and Institute of

Environmental Research (INFU), FR Germany, for

instrument donations.

Nomenclature

A-acetic acid

AI-atomic mass of iodine

AO-atomic mass of oxygen

AAD-average absolute deviation

a-interphase area (m2/m3)

+3H O

C -concentration of hydronium ions (mol/L)

Cj-concentration of component j (mol/L)

Cs-concentration of catalytically active sites (mol/g

catalyst)

D-double bond

E-epoxy group, i.e., ring

EO-epoxy oxygen content (wt.%)

EOESO-epoxy oxygen content in epoxidized soybean

oil (wt.%)

ESO-epoxidized soybean oil

F-objective function

Fj- molar flow of component j (mol/min)

H-hydrogen peroxide

HA-hydroxy acetate

IN-iodine number (g iodine/100 g oil)

IN0-initial iodine number (g iodine/100 g oil)

[j]-amount of component j (mol/100 g oil)

K1-chemical equilibrium constant for peracetic acid

formation

Kj-partition coefficient for component j

KA,SO-partition coefficient for acetic acid in the system

soybean oil-acetic acid-water

KA,ESO-partition coefficient for acetic acid in the system

epoxidized soybean oil-acetic acid-water

ki-rate coefficient for reaction i

ki,0-constant of reparameterized Arrhenius equation

related to the frequency coefficient

a,Eik -constant of reparameterized Arrhenius equation

related to the activation energy

kL,j-mass transfer coefficient for component j (m/min)

Mj-molecular mass of component j (g/mol)

m-mass of the catalyst (g)

mHS-mass of the hydrogen peroxide aqueous solution

(g)

Nj-number of moles of component j (mol)

NR-total number of reactions

NRN-number of runs

NSk-number of samples in run k

P-peracetic acid

R-universal gas constant (J/mol·K)

ri-rate of reaction i

REY-relative epoxy yield (%)

S-stirring speed (rpm)

SE-selectivity

SO-soybean oil

T-temperature (K)

Ta-average temperature of runs (K)

t-reaction time (min)

tHS-period of hydrogen peroxide aqueous solution

addition (min)

tTI-period of temperature increase (min)

V-volume (L)

v-volume (L/100 g oil)

W-water

wH-mass fraction of hydrogen peroxide in its aqueous

solution

X-conversion of double bond (%)


110

Greek letters

αj,i-stoichiometric coefficient of component j in the

reaction i

j -density of component j (g/L)

oo -density of oil in the oil phase (g/L)

SO -density of soybean oil (g/L)

ESO -density of epoxidized soybean oil (g/L)

Subscript

0-initial value

m-maximum value

th-theoretical value

Superscript

aq-aqueous phase

calc-calculated value

exp-experimentally determined value

n-order of reaction

o-organic phase

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111

MILOVAN JANKOVIĆ

SNEŽANA

SINADINOVIĆ-FIŠER

OLGA GOVEDARICA

JELENA PAVLIČEVIĆ

JAROSLAVA

BUDINSKI-SIMENDIĆ

Univerzitet u Novom Sadu, Tehnološki

fakultet Novi Sad, Bulevar cara Lazara

1, 21000 Novi Sad, Srbija

NAUČNI RAD

KINETIKA EPOKSIDOVANJA SOJINOG ULJA PERSIRĆETNOM KISELINOM FORMIRANOM IN SITU U PRISUSTVU JONOIZMENJIVAČKE SMOLE: PSEUDO-HOMOGENI MODEL

Predložen je kinetički model epoksidovanja biljnih ulja persirćetnom kiselinom formiranom

in situ iz sirćetne kiseline i vodonik-peroksida u prisustvu jonoizmenjivačke smole kao

katalizatora. Model je pseudo-homogen u odnosu na katalizator. Pored osnovnih reakcija

formiranja persirćetne kiseline i epoksidne grupe, model opisuje i sporednu reakciju otva-

ranja epoksidne grupe sa sirćetnom kiselinom. U modelu se razmatraju i raspodela sir-

ćetne i persirćetne kiseline između vodene i organske faze sistema i promena zapremina

faza tokom odvijanja procesa epoksidovanja. Temperaturna zavisnost prividnih koeficije-

nata brzina reakcija je opisana reparametrizovanom Arrhenius jednačinom. Kinetički para-

metri predloženog modela su izračunati fitovanjem eksperimentalnih podataka dobijenih

tokom epoksidovanja sojinog ulja izvođenih pri definisanim reakcionim uslovima. Najveći

prinos epoksida od 87,73% je postignut pri temperaturi od 338 K kada je molski odnos

nezasićenost ulja:sirćetna kiselina:vodonik-peroksid iznosio 1:0.5:1.35 i kada je prime-

njena količina katalizatora Amberlite IR-120H bila 4,04 mas.% u odnosu na ulje. U pore-

đenju sa publikovanim pseudo-homogenim modelima, model predložen u ovom radu bolje

koreliše promene sadržaja dvostrukih veza i epoksidnih grupa tokom procesa epoksido-

vanja.

Ključne reči: sojino ulje, epoksidovanje, persirćetna kiselina, jonoizmenjivačka

smola, kinetika.






113

SALAH H. ALJBOUR1

SULTAN A. TARAWNEH2

ADNAN M. AL-HARAHSHEH1

1Department of Chemical

Engineering, College of

Engineering, Mutah University,

Karak, Jordan 2Department of Civil &

Environmental Engineering,

College of Engineering, Mutah

University, Karak, Jordan

SCIENTIFIC PAPER

UDC 669.1.054.8:666.94:62

https://doi.org/10.2298/CICEQ151002016A

EVALUATION OF THE USE OF STEELMAKING SLAG AS AN AGGREGATE IN CONCRETE MIX: A FACTORIAL DESIGN APPROACH

Article Highlights

• A factorial design methodology was applied to evaluate concrete mix production

• Steel slag was evaluated as an aggregate in concrete mix production

• Influential factors on the compressive strength were proposed

• Possible factor-interaction effects were examined

Abstract

Slag is investigated towards its potential use as an aggregate in concrete mix

production. Full factorial design methodology is applied to study the effect of two

process input variables, namely: slag as coarse aggregate and slag as medium

aggregate on the properties of concrete mix. Additionally, the interaction

between input variables is also examined. Incorporating steel slag aggregate in

the concrete mix affected its compressive strength. Enhanced compressive

strength concrete mix was obtained with 70 wt.% coarse slag aggregate and 70

wt.% medium slag aggregate. Under these proportions, the 28-days compres-

sive strength was higher than the 28-days compressive strength of a concrete

mix prepared from normal aggregate. Strong interaction effect exists between

slag aggregate size on the compressive strength at 7-days curing. Lower com-

pressive strength for the concrete mix might be obtained if improper proportions

of mixed medium and coarse slag aggregate were employed.

Keywords: steelmaking slag; concrete; factorial design; compressive strength.

The aggregates typically account for about 75%

of the concrete volume and play a substantial role in

different concrete properties such as workability,

strength, dimensional stability and durability. Conven-

tional concrete consists of sand as fine aggregate and

gravel, limestone or granite in various sizes and

shapes as coarse aggregate. There is a growing inte-

rest in using waste materials as alternative aggregate

materials and significant research is made on the use

of many different materials as aggregate substitutes

such as coal ash, blast furnace slag, and steel slag

aggregate. This type of use of a waste material can

solve problems of lack of aggregate in various con-

struction sites and reduce environmental problems

Correspondence: S.H. Aljbour, Department of Chemical Engine-

ering, College of Engineering, Mutah University, Karak, Jordan. E-mail: [email protected] Paper received: 2 October, 2015 Paper revised: 12 February, 2016 Paper accepted: 17 March, 2016

related to aggregate mining and waste disposal. The

use of waste aggregates can also reduce the cost of

the concrete production [1].

Jordan’s steel and iron industry began with the

establishment of the first local steel manufacturer,

Jordan Iron and Steel in 1965. Due to continuous and

rising need for steel, twelve local steel mills are now

established in Jordan with a total annual production

capacity of 1.2 million t [2]. Some of these factories

utilize imported semi processed iron in the form of

plates to be melted and processed in a later stage to

produce reinforced bars. The other type of factories

utilizes iron scrap for the production of reinforced

bars. It is estimated that the annual production of slag

from these factories is around 100-200 thousand

tons. Inventing new ways to reuse this accumulated

waste is the most pressing and daunting challenge

that face Jordan’s industrial sector.

The primary components of iron and steel slag

are limestone and some other materials in oxide form.

S.H. ALJBOUR et al.: EVALUATION OF THE USE OF STEELMAKING SLAG… Chem. Ind. Chem. Eng. Q. 23 (1) 113119 (2017)

114

In the case of steelmaking slag, the slag contains

metallic elements such as iron in oxide form; how-

ever, because refining time is short and the amount of

limestone contained is large, a portion of the lime-

stone auxiliary material may remain un-dissolved as

free CaO [3]. In general, several factors are affecting

physical and chemical properties of steel slag. These

factors include type of steel furnace in the steel-

making plant and the method of steel slag processing.

Various approaches can be followed for the utilization

of steel slag in cement and concrete applications.

One may think about using steel slag in the product-

ion of cement. In this approach, the steel slag is

mixed with limestone and clay as a raw material feed

to cement kiln. In this case, the slag must be clinkered

[4]. Another approach is the incorporation of steel

slag in cement and composite cements [5]. In addition

steel slag can be utilized as an aggregate material.

Several advantages of using slag aggregates in con-

crete mixes are gained such as: reliable quality, inc-

reased strength and does not generate alkali-aggre-

gate reactions. In addition, blast furnace slag fine

aggregate does not contain materials that may affect

the strength and durability of concrete, such as chlo-

rides, organic impurities clay and shells [5,6].

Extensive research has been conducted for the

application of steel slag in broad areas of construct-

ion. It is vital to quantify the benefits of using such

cheap material in concrete technology and concrete

asphalt pavement [6-14]. In this study, we evaluate

the beneficial use of steel slag obtained from Jordan’s

steel industry as an aggregate in concrete mix pro-

duction. Our investigation involves studying the effect

of using steel slag of different grades when combined

with normal aggregates by different ratios on impro-

ving the mechanical properties of hardened concrete.

The effect of medium slag aggregate and coarse slag

aggregate and the combination between them on the

mechanical quality of concrete mix is investigated by

applying a 22 full factorial design methodology. The

main objectives of this study are to identify the most

influential process operating conditions on the pro-

duction of concrete mix and the interaction effects

among variables on the compressive strength of the

produced concrete.

EXPERIMENTAL

Preparation of the steel slag-based aggregate

Steel slag was provided by local Jordanian fac-

tory in the form of boulders (size < 100 mm). The

factory produces steel products made from scrap

metal, recycled from used automobiles, plant equip-

ment, machinery, or byproducts of the manufacturing

and construction sectors by utilizing an electric arc

furnace (EAF) for scrap melting and a ladle furnace

for the precision control of chemistry and the puri-

fication of the liquid steel.

The majority of the steel slag contains free CaO

and MgO. Experiments must be performed to inves-

tigate the content of free CaO and MgO in the slag. In

general, the content of free CaO and MgO in EAF-

-slag is significantly lower than in basic oxygen fur-

nace steel slag (BOF slag) [1]. Slag pretreatment is

necessary to reduce the content of free CaO and

MgO in the slag. The ageing or weathering method

was followed to reduce the content of free CaO and

MgO in the EAF slag. It has been reported that that a

proper treatment aimed to stabilize slag by exposing

them to outdoor weather and regular spraying for at

least 90 days, may eliminate any subsequent exp-

ansive phenomenon, allowing a safe use of such slag

as aggregate in concrete production [15]. Prior to use,

the EAF slag was aged for a period of 6 months. The

air aging method was applied by leaving the EAF slag

out in an open area to enable weathering. The pre-

sence of free CaO and MgO in the slag does not

seem to represent a limit for the durability of concrete,

due to their stabilization in crystalline lattice [16].

The aged EAF slag aggregate was prepared by

crushing the boulders, followed by sorting the ground

slag by sieving. In this study, medium and coarse

EAF slag was used. The medium EAF slag aggregate

was obtained from the sieved material which passed

through the 12.5 mm sieve and retained on the 4.75

mm sieve. The coarse EAF slag aggregate was

obtained from the sieved material which passed

through the 19 mm sieve and retained on the 12.5

mm sieve. Physical and chemical properties of the

EAF slag aggregates are given in Table 1. The mech-

anical properties of the EAF slag aggregates were

conducted according to (ASTM C 33, 2003; ASTM C

138/C 138M, 2001; ASTM C 150, 2005).

Table 1. Mechanical properties of aggregate used in this study

Parameter Steel slag

aggregate

Natural

aggregate

Saturated density, kg/m3 2323 2020-2100

Apparent specific gravity 3.2 -

Abrasion value, % 19.4 20-24

Flakiness index 10.98 20-30

Elongation index 9.89 10

Crushing value, % 26.1 -


115

Preparation of normal aggregate

The normal coarse, medium and fine aggregate

source for all mix designs was obtained from gravel

pits. The gravel was obtained from local area known

in Jordan as “Al-Ghoor”. The gravel was crushed then

sieved to obtain the desired size fraction. The fine

aggregate was obtained from sieved material which

passed through the 4.75 mm. The medium aggregate

was obtained from the sieved material which passed

through the 12.5 mm sieve and retained on the 4.75

mm sieve. The coarse aggregate was obtained from

the sieved material which passed through the 19 mm

sieve and retained on the 12.5 mm sieve.

Preparation of concrete mix

All concrete mixes were prepared by keeping

the water-to-cement ratio and fine aggregate-to-total

aggregate ratio constant at 0.6 and 0.35, respectively.

No additives were used for the preparation of con-

crete mix. The high water-to-cement ratio was applied

to ensure good workability conditions since no addi-

tives were used for concrete mix preparation and to

account for water absorption by aggregates. Cement

used in this study was Ordinary Portland Cement pro-

duced by a local Jordanian factory. The cement is

classified as CEM-1 42.50 N. The chemical compo-

sition of cement is as follows: 19.94% SiO3, 5.37%

Al2O3, 3.18% Fe2O3, 63.65% CaO, 2.59% MgO,

2.88% SO3, 0.82% K2O and 0.1% Na2O [17].

The materials were added to the concrete mix-

ture in the following order: coarse aggregate, medium

aggregate, fine aggregate and cement. The mixture

was mixed under dry condition for about 1 min, then

80 % of water was added. After 1.5 min of mixing, the

rest of water was added. Every batch of concrete mix

was mixed for a total time of 3 min. The concrete

mixes were cast in steel molds (150 mm150

mm150 mm) and compacted using a tamping rod.

One day after casting, the concrete samples were

removed from the mold and cured in a tank of water

at a temperature of 20 C for 3, 7 and 28 days. All

concrete mix are prepared by keeping the cement

content as 425 kg/m3, medium-to-total aggregate ratio

of 0.38 and coarse-to-total aggregate ratio of 0.6.

Table 2 shows the mix proportions for the mixes

applied in this study.

Characterization of samples

Sample characterization was carried out by

examining the compressive strength. The compres-

sive strength of concrete was determined according

to ASTM C-39.

Experimental factorial design and analysis

Design of experiment is a powerful tool that can

be used in a wide spectrum of experimental situ-

ations. Design of experiments allows for multiple input

factors to be studied and to determine their effect on a

desired process/design/quality output. When studying

multiple inputs at the same time, design of experiment

can identify important interactions that may missed

when experimenting with one variable at a time

(OVAT approach). All possible combinations between

process input variables can be investigated by con-

ducting full factorial design. The factorial design

methodology can be utilized to confirm possible input/

/output relationships and to develop a predictive

equation suitable for performing design simulations

with minimum time and cost.

In this research, EAF slag is incorporated as

aggregate during the production of concrete mix.

Most of studies concerning concrete containing slag

aggregate are conducted by adopting the OVAT

approach. In these studies, a given grade of con-

ventional aggregate are replaced partially or totally by

its counterpart slag aggregate with the remaining

factor held constants. The effect of such replacement

of either fine, medium or coarse aggregate is then

solely evaluated. This approach provides an estimate

of the effect of a single variable at a selected fixed

condition of the other variables. However, for such an

estimate to have general relevance it is necessary to

check whether the effect would be the same at other

settings of the other variables or not. Nevertheless,

studies on the effect of combined replacement of

Table 2. Mix design parameters for concrete mixes applied in this study (kg/m3)

Component Run-1 Run-2 Run-3 Run-4 Run-5

Cement 425 425 425 425 425

Water 255 255 255 255 255

Fine aggregate 510 510 510 510 510

Medium slag aggregate 138 415 138 415 0

Medium normal aggregate 415 138 415 138 553

Coarse slag aggregate 223 223 668 668 0

Coarse normal aggregate 668 668 223 223 891


116

mixed grade of aggregate on the mechanical pro-

perties of concrete at the same time are rarely con-

ducted.

To achieve this goal, a full factorial design meth-

odology was followed to identify the main effects of

two processing factors on the mechanical properties

of concrete. The two factors studied were replace-

ment percentage of medium conventional aggregate

by medium EAF slag aggregate (X1), and replace-

ment percentage of coarse conventional aggregate by

coarse EAF slag aggregate (X2). Each of the two

factors was studied at two levels (Table 3). Therefore,

the arrangement and number of experiments is

considered to be 2x2 or 22 factorial design. The four

formulations are shown in Table 4 with variable levels

coded with plus and minus signs. The prepared con-

crete mixes were subjected to the following tests (res-

ponse output variables): 3-days compressive strength

(Y1), 7-days compressive strength (Y2), and 28-days

compressive strength (Y3). In addition, one concrete

mix that contained 100% normal aggregate was pre-

pared and tested for its 3, 7 and 28-days compressive

strength.

Table 3. Process input variables and levels of variables

Input variable Slag as medium

aggregate, wt.% (X1)

Slag as a coarse

aggregate, wt.% (X2)

Level Low (-) High (+) Low (-) High (+)

Condition 30 70 30 70

The main effect of each variable is computed

using the following equation [18]:

1 1E Y Y

where 1Y is average response for the high level of

the variable and 1Y is average response for the low

level of the variable.

The interaction ( A,BI ) between two process

parameters (say, A and B) can be computed using the

following equation [18,19]:

A,B A,B( 1) A,B( 1)

1

2I E E

where A,B( 1)E is the effect of factor “A” at high level of

factor “B” while A,B( 1)E is the effect of factor “A” at low

level of factor “B”.


Main and interaction effects of process input variables

on compressive strength

The main effects of process variables on the 3, 7

and 28 day compressive strength have been studied.

Table 5 shows the calculated main effects based on

all experimental observations. Figure 1 shows the

main effect plots for the studied process variables.

The absolute value of the effect of EAF slag

aggregate grade on the 3-days compressive strength

was the same. Very slight effect was noticed for both

aggregate grades at any proportion on the 3-days

compressive strength. Concrete mix made of either

30 or 70 wt.% coarse EAF slag aggregate and con-

crete mix made of either 30 or 70 wt.% medium EAF

slag aggregate exhibited almost the same (21.9–23.4

MPa) 3-days compressive strength. In addition, con-

crete mix made from normal aggregate without incor-

porating any of the steel slag aggregate possessed

nearly the same 3-days compressive strength of 22.8

MPa. It seems that three days of curing was not

sufficient enough to produce strong adhesion between

the cement paste and the aggregate, thereby, giving

close value of compressive strength. To assess the

viability of effect of parameters and confirm the exist-

ence or absence of interaction effects between para-

meters, we constructed the interaction plot shown in

Figure 2.

Table 5. Full 22 factorial design analysis of process response

Term Response

Y1 Y1 Y3

Main Effect of X1 -1.6 -2.2 4.6

Main Effect of X2 -1.6 3.7 6.3

Interaction effect X1X2 0.5 3.7 -1.9

Figure 2 shows that there is week interaction

effect between parameters, that is, the main effect of

Table 4. Experimental design matrix and response of output variables

Run In put variables Output variables

X1 X2 Y1 / MPa Y2 / MPa Y3 / MPa Slump, cm

1 - - 24.4 31.1 34.6 0

2 + - 22.4 25.3 41.1 2

3 - + 22.4 31.1 42.8 1

4 + + 21.3 32.6 45.5 2

5 0% 0% 22.8 31.1 40.9 0


117

coarse EAF slag aggregate proportion in the concrete

mix on the 3-days compressive strength is indepen-

dent on the proportion of medium EAF slag aggregate

in the same mix. Similarly, the main effect of medium

EAF slag aggregate proportion in the concrete mix on

the 3-days compressive strength is independent on

the proportion of coarse EAF slag aggregate in the

same mix.

Figure 2. Interaction effect plot for coarse and medium slag

aggregate on 3-days compressive strength.

The 7-days curing enhanced the compressive

strength of concrete mixes. The 7-days compressive

strength of concrete mixes containing EAF slag

aggregate was larger than 3-days compressive

strength by at least 30%. The same results was

noticed for concrete mix made from normal aggre-

gate without incorporating any of the EAF slag aggre-

gates. The effect of medium EAF slag aggregate on

the 7-days compressive strength was slightly higher

at 30 wt.% replacement compared to 70 wt.%

replacement. On the other hand, the effect of coarse

EAF slag aggregate on the 7-days compressive

strength was slightly higher at 70 wt.% replacement

compared to 30 wt.% replacement. This trend in

results must also be judged by confirming existence

or absence of interaction effect between EAF slag

aggregate grades. This is shown in the interaction

plot given in Figure 3.



The interaction between the process input vari-

ables on the 7-days of compressive strength repre-

sents an antagonistic-type interaction [19]. At 75 wt.%

medium slag proportion, increasing the coarse EAF

slag aggregate proportion in the concrete mix from 30

to 70 wt.% substantially increased the 7-days com-

pressive strength from 25.3 to 32.6 MPa. However, at

30 wt.% medium EAF slag proportion increasing the

coarse EAF slag aggregate proportion in the concrete

mix from 30 to 70 wt.% kept the 7-days compressive

strength unchanged at 31 MPa. This interaction

effect between parameters can hardly be observed by

following the OVAT approach of experimentation. The

interaction effect shown in Figure 3 shows that local

maximum of 7-days compressive strength could be

obtained at 70 wt.% medium EAF slag aggregate

proportions and 70 wt.% coarse EAF slag aggregate

proportions in the concrete mix.

Figure 1. Main effect plots for process input variables (coarse and medium slag aggregate) on compressive strength.


118

The 28-days curing enhanced the compressive

strength of concrete mixes. The 28-days compressive

strength of concrete mixes containing EAF slag

aggregates was larger than 3-days compressive

strength by at least 62%. Concrete mix made from

normal aggregate exhibited 40.9 MPa of 28-days

compressive strength with an enhancement by appro-

ximately 80% in reference to the 3-days compressive

strength of the concrete mix made from normal aggre-

gate. The effect of medium EAF slag aggregate on

the 28-days compressive strength was higher at 70

wt.% replacement compared to 30 wt.% replacement.

Similarly, the effect of coarse EAF slag aggregate on

the 28-days compressive strength was noticeably

higher at 70 wt.% replacement compared to 30 wt.%

replacement. This trend in results must also be jud-

ged by confirming existence or absence of interaction

effect between slag aggregate grades. This is shown

in the interaction plot given in Figure 4.



Small interaction between the process input

variables on the 28-days of compressive strength was

noticed within the concrete mix proportions selected

in this study. At 70 wt.% medium EAF slag proportion,

increasing the coarse EAF slag aggregate proportion

in the concrete mix from 30 to 70 wt.% increased the

28-days compressive strength from 41.1 to 45.5 MPa.

In the same manner, at 30 wt.% medium EAF slag

proportion increasing the coarse EAF slag aggregate

proportion in the concrete mix from 30 to 70 wt.%

increased the 28-days compressive strength from 34.6

to 42.8 MPa. The interaction effect shown in Figure 4

shows that local maximum of 28-days compressive

strength could be obtained at 70 wt.% medium EAF

slag aggregate proportion and 70 wt.% coarse EAF

slag aggregate proportion in the concrete mix.

CONCLUSION

Concrete mixes were successfully prepared by

utilizing slag generated from the steelmaking industry.

A full factorial design analysis was effectively per-

formed to assess the most influential process oper-

ating conditions on the compressive strength of con-

crete mixes made from different mixed proportions of

EAF slag aggregates. The proportions of the course

and medium size EAF slag aggregate were prominent

process variables found to be affecting the 7-days

and 28-days compressive strength of a concrete mix.

Under 3-days curing, there was no appreciable

enhancement on the compressive strength of con-

crete mixes prepared from EAF slag aggregate at any

proportion in comparison with concrete mixes prepared

from normal aggregates. Both type of mixes pos-

sessed almost the same 3-days compressive strength.

Under 7-days curing, strong interaction effect for

the proportion of EAF slag aggregates was noticed on

the compressive strength. Concrete mixes comprising

lower coarse slag aggregate content and higher

medium EAF slag content exhibited lower compres-

sive strength compared with concrete mix prepared

from normal aggregate. Nevertheless, concrete mixes

comprising higher coarse slag aggregate content and

any proportion of medium EAF slag content exhibited

almost the same compressive strength of concrete

mix prepared from normal aggregate.

Under 28-days curing, there was noticeable

enhancement on the compressive strength of con-

crete mixes prepared from high proportions of EAF

slag aggregate in comparison with concrete mixes

prepared from normal aggregates.

The mechanical properties of concrete mixes

prepared in this study are satisfactory. However, it is

recommended that other important properties such as

durability and corrosion of EAF slag concrete are to

be investigated before mass use.

Acknowledgement

Lab engineers: Wafaa Suhaimat and Hussein

Sarayreh of the civil engineering department and Ali

Alzoubi of the Chemical Engineering Department at

Mutah University are acknowledged for laboratory

assistance.

REFERENCES

[1] J. de Brito, N. Saikia, Recycled Aggregate in Concrete,

Use of Industrial, Construction and Demolition Waste,

Springer-Verlag, London, 2013, pp. 23,34

[2] Jordan Steel Company, Global Research, Global Invest-

ment House, Kuwait, Jan 2009


119

[3] Chemical characteristics of iron and steel slag, Nippon

Slag Association, http://www.slg.jp/e/slag/character.html

(accessed 1 Oct, 2015)

[4] A. Monshia, M.K. Asgarani, Cem. Concr. Res. 29 (1999)

1373-1377

[5] S. Kourounis, S. Tsivilis, P.E. Tsakiridis, G.D. Papadi-

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[6] S.A. Tarawneh., E.S. Gharaibeh, F.M. Saraireh, Am. J.

Appl. Sci. 11 (2014) 700-706

[7] F.H. Al-Negheismish, A.I. Al-Sugair, R.Z. Al-Zaid, J. King

Saud Univ., Eng. Sci. 9 (1996) 39-55

[8] M. Maslehuddin, A.M. Sharif, M. Shameem, M. Ibrahim,

M.S. Barry, Constr. Build. Mater. 17 (2003) 105-112

[9] H.Y. Moon, J.H. Yoo, J. Korea Concr. Inst. 11 (1999) 101-

–111

[10] P.T. Sherwood, Alternative Materials in Road Cons-

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Aggregates, Thomas Telford Publishing, London, 2001,

p. 97

[11] I.M. Asi, H.Y. Qasrawi, F.L. Shalabi, Can. J. Civ. Eng. 34

(2007) 902-911

[12] J.M. Manso, J.J. Gonzalez, J.A. Polanco, J. Mater. Civ.

Eng. 16 (2004) 639-645

[13] H. Beshr, A.A. Almusallam, M. Maslehuddin, Constr.

Build. Mater. 17 (2003) 97-103

[14] R. Alizadeh, M. Chini, P. Ghods, M. Hoseini, Sh. Mon-

tazer, M. Shekarchi, Utilization of Electric Arc Furnace

Slag as Aggregates in Concrete - Environmental Issue, in

Proceedings of the 6th CANMET/ACI International Con-

ference on Recent Advances in Concrete Technology,

Bucharest, Romania (June 2003), p. 451

[15] C. Pellegrino, V. Gaddo, Cem. Concr. Compos. 31 (2009)

663-671

[16] C. Pellegrino, P. Cavagnis, F. Faleschini, K. Brunelli,

Cem. Concr. Compos. 37 (2013) 232-240

[17] H. Al-Hamaideh, O. Maaitah, S. Mahadin, Electron. J.

Geotech. Eng. 15 (2010) 601-608

[18] G.E.P. Box, W.G. Hunter, J.S. Hunter, Statistics for Expe-

rimenters –Design, Innovation and Discovery, John Wiley

and Sons Inc., Hoboken, NJ, 2005, p. 180

[19] J. Antony, Design of Experiments for Engineers and Sci-

entists. Elsevier Science & Technology, Burlington, 2003,

p. 19.

SALAH H. ALJBOUR1

SULTAN A. TARAWNEH2

ADNAN M. AL-HARAHSHEH1

1Department of Chemical Engineering,

College of Engineering, Mutah

University, Karak, Jordan 2Department of Civil & Environmental

Engineering, College of Engineering,

Mutah University, Karak, Jordan

NAUČNI RAD

PROCENA UPOTREBE ŠLJAKE IZ PROIZVODNJE ČELIKA KAO AGREGATA U CEMENTNIM MEŠAVINAMA POMOĆU PUNOG FAKTORIJALNOG PLANA

U radu je analizirana upotreba šljake kao potencijalnog agregata u proizvodnji cementa.

Primenjen je pun faktorijalni plan radi proučavanja uticaja dve ulazne procesne promen-

ljive, i to šljaka kao krupni agregat i šljaka kao srednji agregat na osobine cementne meša-

vine. Pored toga, ispitivana je interakcija između ulaznih promeljivih. Ubacivanje agregata

šljake iz proizvodnje čelika u cementnu mešavinu uticalo je na njenu pritisnu čvrstoću.

Povećana pritisne čvrstoće cementne mešavine sa 70 mas.% krupne šljake i sa 70 mas.%

srednje šljake. Pri ovim proporcijama, 28-dnevna pritisna čvrstoća dobijene cementne

mešavine je jača od 28-dnevne pritisne čvrstoće cementne mešavine pripremljene od nor-

malnih agregata. Postoji jaka interakcija između veličine agregata šljake i pritisne čvrstoće

posle sedmodnevnog očvršćavanja. Manja pritisna čvrstoća cementne mešavine mogla bi

biti dobijena ako bi se koristile agregati srednje i krupne šljake pomešani u neodgovara-

jućim proporcijama.

Ključne reči: bioetanol, pamuk, lignoceluloza, predtretman, hidroliza, fermentacija.






121

DUŠAN Lj. PETKOVIĆ

MILOŠ J. MADIĆ

GORAN M. RADENKOVIĆ

University of Niš, Faculty of Mechanical

Engineering, Niš, Serbia

SCIENTIFIC PAPER

UDC 669.14.018.8:546.175–323:51

https://doi.org/10.2298/CICEQ151127020P

THE EFFECTS OF PASSIVATION PARAMETERS ON PITTING POTENTIAL OF BIOMEDICAL STAINLESS STEEL

Article Highlights

• Corrosion resistance of 316LVM stainless steel was increased by passivation

• Multiple regression analysis and artificial neural network (ANN) were employed

• Only the ANN model provided a statistically accurate mathematical model

• Pitting potential is highly non-linearly dependent on the passivation parameters

• Nitric acid concentration has the strongest influence on the pitting potential

Abstract

Passivation is a chemical process in which the electrochemical condition of

passivity is gained on the surface of metal alloys. Biomedical AISI 316LVM

stainless steel (SS) can be passivized by means of nitric acid immersion in order

to improve a protective oxide layer on the surface and consequently increase

corrosion resistance of the SS in the physiological solutions. In this study, mul-

tiple regression analysis and artificial neural network (ANN) were employed for

mathematical modeling of the AISI 316LVM SS passivation process after immer-

sion in the nitric acid solution. The pitting potential, which represents the mea-

sure of pitting corrosion resistance, was chosen as the response, while the pas-

sivation parameters were nitric acid concentration, temperature and passivation

time. The comparison between experimental results and models predictions

showed that only the ANN model provided statistically accurate predictions with

a high coefficient of determination and a low mean relative error. Finally, based

on the derived ANN equation, the effects of the passivation parameters on pitting

potential were examined.

Keywords: stainless steel, nitric acid, passivation, multiple regression analysis, artificial neural networks.

AISI 316LVM is a vacuum melted stainless steel

(SS) widely used for biomedical applications. It has

high tensile strength and fatigue resistance, good

deformability, and relatively low price. Examples of its

biomedical applications include bone plates and

screws, hip and knee prosthesis, nails and pins, den-

tal prostheses as well as vascular and urological

stents [1]. The main disadvantages of this steel are

local corrosion susceptibility during prolonged contact

with human tissue, and release of metal ions [2].

Additionally, nickel is known as a strong immuno-

logical reaction medium and may cause various

Correspondence: D.Lj. Petković, University of Niš, Faculty of

Mechanical Engineering, A. Medvedeva 14, 18000 Niš, Serbia. E-mail: [email protected] Paper received: 27 November, 2015 Paper revised: 2 April, 2016 Paper accepted: 6 April, 2016

health problems [3]. Despite the above listed weak-

nesses, SS has the ability to spontaneously form a

stable self-protecting oxide layer (passive film) on its

surface in the reaction with air or most aqueous envi-

ronments. This film consists mostly of chromium

oxide (Cr2O3) and typically shows thicknesses of few

nanometers [4]. The presence of nonmetallic inc-

lusions on the material’s surface, such as sulfide inc-

lusions, represents a discontinuity of the passive film

and therefore a potential place of pitting corrosion

initiation [5].

Localized corrosion may cause an accidental

deterioration of the whole system with disastrous con-

sequences, while the total mass loss is insignificant

[6]. Corrosion of SS implants have two effects [7]:

first, the implant may become weak and the pre-

mature failure of the implant may happen; and sec-

D.Lj. PETKOVIĆ et al.: THE EFFECTS OF PASSIVATION PARAMETERS… Chem. Ind. Chem. Eng. Q. 23 (1) 121129 (2017)

122

ond, the release of corrosion products from the imp-

lant can cause the tissue reaction.

Corrosion pits commonly start because of chem-

ical or physical heterogeneities at the surface, which

include dislocations, mechanical damage, inclusions,

or second phase particles [8]. The resistance of SS to

a pitting attack depends largely on the type of SS

used, and on the subsequent physicochemical pro-

perties of the protective passive oxide layer formed

on its surface [9]. There has been a constant attempt

by engineers and scientists to improve the surface-

related properties of biomedical materials to reduce

the failure of implants and leaching of ions due to

wear and corrosion. A number of research groups

have done extensive research on the improvement of

both general and pitting corrosion resistance of SS by

developing techniques for the modification of the mat-

erial’s surface and passive film. Further, pitting attack

resistance directly depends on the physicochemical

properties of the protective passive oxide layer formed

on the surface [10].

A beneficial effect of nitric acid solution on chro-

mium enrichment in the modified passive layer of SS

was reported in the literature [11-14]. Immersion in

nitric acid solutions is particularly effective in impro-

ving the pitting resistance of austenitic SS [15]. Also,

immersion in nitric acid removes sulphide inclusions,

eliminating the preferential sites for attack [16].

Mathematical modeling of the passivation pro-

cess based on the scientific principles allows one to

study and better understand this complex process.

Multiple regression analysis (MRA) and ANNs are two

important competitive data mining techniques widely

used for the development of predictive mathematical

models [17]. MRA is a conceptually simple method for

development of the functional relationships between

several independent (input) variables and one depen-

dent (output). ANNs are a computational tool, based

on the properties of biological neural systems, which

have been used successfully where conventional

computer systems have traditionally been slow and

inefficient. Both methodologies can be successfully

applied for different process modeling. However,

when compared to one another, different conclusions

can be drawn in certain cases [18].

In the literature there are few studies which are

aimed at modeling the passivation process of biomed-

ical material in nitric acid as well as in other fluids in

general. Masmoudi et al. [19] studied the passivation

process of CP Ti (commercially pure titanium) and

Ti6Al4V alloy by immersing in HNO3 solution. Their

main aim was to improve corrosion resistance of

tested materials after acid treatment. Mathematical

models were obtained by employing MRA, while opti-

mization was performed by applying the least square

method. Jiménez-Come et al. [20] presented an auto-

matic model based on artificial intelligence techniques

to predict pitting potential values. Their model was

aimed to compare pitting corrosion resistance of AISI

316L austenitic SS in different environmental condi-

tions without requiring the use of electrochemical tests.

They showed that the presented model provides an

automatic way to compare the pitting corrosion resist-

ance of austenitic stainless steel in different environ-

mental conditions. Petković et al. [13] analyzed the

possibilities for enhancing corrosion resistance of bio-

medical AISI 316LVM SS by immersing in nitric acid

solutions under different passivation conditions.

Namely, the effects of nitric acid solution concentra-

tion, temperature and passivation time on the pitting

potential, which was selected as a parameter for cor-

rosion resistance assessment, were investigated. A

total of 27 experimental trials were carried out

according to 33 full experimental design. A mathemat-

ical model was determined by using MRA, while opti-

mal values for passivation parameters were found by

means of genetic algorithm.

We decided to broaden the previous research in

order to obtain more precise and accurate experimen-

tal results by repeating the experiment twice more

and by applying ANNs for the purpose of mathemat-

ical modeling. Thus, in this paper, the pitting potential

value, as the response, was calculated as a mean of

the three pitting potential values measured for all 27

experiment trials (test). Moreover, three additional

measurements were carried out in order to determine

the pitting potential for non-passivized sample (con-

trol). Hence, there were a total of 84 experiment trials.

Compared to our previous study, the application of

the experiment designs with replications increases its

reliability significantly.

The aim of this study was to develop a mathe-

matical model relating passivation parameters with

pitting potential as the response. To evaluate the best

possible mathematical model, a statistical analysis of

the results was performed. Based on the conducted

statistical analysis, one can argue that MRA is not

able, on a satisfactory level, to accurately model the

underlying relationships between passivation para-

meters and pitting potential. For this reason, a mathe-

matical model of the passivation process was deve-

loped by using ANN in combination with a 33 full fac-

torial design with three replications. Finally, the ANN

model was compared with the MRA model to assess

the adequate methodology for further modeling of the

similar processes.


123

EXPERIMENTAL

Biomedical SS

For this research, 81 test samples (3 per each of

27 experimental trials) and 3 control samples were

machined. The samples were cylindrical with the dia-

meter of 6 mm and height of 20 mm made of AISI

316LVM SS, containing Cr, Ni, Mo and Mn as main

alloy elements. Chemical composition of tested AISI

316LVM SS is in accordance to ISO 5832-1 [21].

Passivation process

Three input variables (X1: HNO3 concentration,

X2: temperature of passivation solution, and X3: passi-

vation time) were selected as passivation parameters.

The 33 full factorial design with three replications was

used. Real and coded values of the parameters and

their levels used in the experimentation are given in

previous published paper [13].

Prior to each test, the exposed surface of the

samples was wet ground with silicon carbide paper up

to 1200 grit and polished by using diamond paste with

grain size of up to 0.25 µm. Then, the samples were

rinsed with distilled water and washed with ethanol in

an ultrasonic cleaner. The passivation treatment was

performed by immersing samples in nitric acid sol-

utions. Lastly, the samples were rinsed in double dis-

tilled water and alcohol, respectively.

Electrochemical measurements

Electrochemical tests for each sample were per-

formed using a three-compartment cylindrical glass

cell equipped with a saturated calomel electrode

(SCE) as the reference electrode and a platinum foil

as the counter electrode. The average of pitting pot-

entials for three samples with the same treatment was

chosen as a measure of corrosion resistance. The

specimens were immersed 15 s before the start of the

potential rise and this time was set by the program.

The starting potential was –400 mV with a scan rate of

0.25 mV/s to anodic potential direction. The tests were

finished when the current density reached about 0.2

mA/cm2. The pitting potential (Ep) was chosen as a

measure of corrosion resistance and represented a

level of potential when the passive film broke down [22].

The electrochemical tests were conducted in

Hank’s solution, which is a simulated body fluid and

most frequently used for in vitro tests. During the

experiments, the temperature was maintained at

37±1 C (typical body temperature). The composition

and instruction for preparation of the Hank’s solution

are described elsewhere [23].

Mathematical models

MRA model

In this study, a second order polynomial was

selected for mathematical modeling of pitting potential

depending on the passivation parameters forms, as

follows:

p 0 1 1 2 2 3 3 12 1 2

2 2 213 1 3 23 2 3 11 1 22 2 33 3

E b b X b X b X b X X

b X X b X X b X b X b X (1)

where pE is the pitting potential (output), jX are

coded values of the parameters (input), 0b is the

model constant, jb is the first degree coefficient, jkb

are the cross-products coefficients and jjb are the

quadratic coefficients.

The regression coefficients, 0b , jb , jkb and

jjb , were estimated by the least squares method.

Values of regression coefficients and their statistical

significance were determined by using Minitab 15 sta-

tistical software package.

ANN model

Three neurons in the input layer (for each of the

passivation parameters), one neuron at the output

layer for pitting potential, and only one hidden layer

were used to define the ANN architecture [18,24]. The

number of hidden neurons was selected by consider-

ing the following: i) too few neurons in the hidden

layer can lead to under-fitting, i.e., inability to perform

appropriate function approximation, whereas too

many neurons can contribute to over-fitting [25],

which results in a lack of generalization capability of

the developed model; ii) the more hidden neurons,

the more expressive power of the ANN – however,

with the increase of the number of hidden neurons,

the number of unknown parameters (weights and

biases) to be estimated also increases; (ii) the upper

limit of the number of hidden neurons can be deter-

mined considering that the total number of unknown

parameters does not exceed the number of available

data for training process. As noted by Sha and

Edwards [26], although in the case where the number

of the connections to be fitted is larger than the num-

ber of available data for training, ANN can still be

trained, the case is mathematically undetermined.

Therefore, relatively small ANN architecture 3-5-1

was selected to model this passivation process.

Since it was assumed that a nonlinear relation-

ship exists between the passivation parameters and

pitting potential, the hyperbolic tangent sigmoid trans-

fer (activation) function was used in the hidden layer,

and linear transfer function was used in the output

layer. According to the selected transfer functions in

the input and output layer, all experimental data were


124

normalized in the [1,1] range. The goal of the ANN

training process is to determine (near) optimal values

of weights and biases in the hidden and output layer,

previously initialized by the Nguyen-Widrow algo-

rithm, in order to minimize the mean squared error

between ANN predictions and experimental data. The

ANN was trained with gradient descent with mom-

entum by using 23 out of 27 sets of input/output expe-

rimental data and the rest was used for testing the

ANN’s generalization performance capability. Learn-

ing rate () and momentum (µ) were kept at 0.1 and

0.9, respectively. The training process was finished

after 7800 epochs, with the minimal achieved mean

squared error of 0.00548.

Statistical evaluation of developed models

Coefficient of determination R2 was used to

evaluate the performance of the developed models,

and indicate how well mathematical models fitted

experimental data [24]. In addition, for the estimation

of the prediction performance of the developed math-

ematical models, relative error, as one of the most

stringent criteria, was calculated by using the follow-

ing equation:

Relative error

Experimental value Predicted value100

Experimental value

(2)

The mean relative error (MRE) was also calcul-

ated.


The results of modeling the passivation process

by using MRA and ANN are displayed and compared

in this section. In addition, the results are discussed

and analyzed.

The second order MRA model (full quadratic

regression model with interactions), relating passiv-

ation parameters and the pitting potential, was

obtained as:

p 1 2 3

2 21 2 2 3 1 2

1.50 0.086 0.0014 0.0334

0.077 0.056 0.227 0.130

E X X X

X X X X X X (3)

Based on the Table 1, it should be noted that

insignificant model terms X1X3 and X32 were elimin-

ated since they were highly correlated with other

variables.

The R2 value indicates that the passivation para-

meters explain 53.7 % of variance in pitting corrosion

potential. Apart from that, adjusted coefficient of det-

ermination and predicted coefficient of determination

(given in Table 1) are considerably smaller indicating

that the model is inadequate and over-fitted. Con-

sequently, the MRA model is not reliable enough to

describe the investigated relationship. Moreover,

analysis of variance (ANOVA), shown in Table 2, rev-

eals that F-ratio of 2.19 corresponds to confidence

level of 92.2%, which is lower than standard (95 or

99%).

Based on the conducted statistical analysis, one

can argue that MRA model is not able, on a satis-

factory level, to accurately model the underlying rel-

ationships between passivation parameters and pit-

ting potential. For these reasons, modeling of the pas-

sivation process was attempted by means of ANN.

By considering the data normalization, transfer

functions used in the hidden and output layer, and by

using the weights and biases from Table 3, one can

obtain a mathematical equation for pitting potential

calculation. After denormalization, the mathematical

model for pitting potential in terms of the passivation

parameters can be expressed by the following equation:

1 1p 2 22

20.36 1 1 0.77

1X W B

E W Be

(4)

Table 1. Regression coefficients of the MRA model; S = 0.174801; R2 = 53.7%; R2(adj.) = 29.2%; R2(pred.) = 0.0%

Coefficient Calculated coefficient value SE Coefficient T Probability density P

b0 1.5003 0.9300 16.13 0.000

b1 0.0856 0.0412 2.08 0.053

b2 0.0014 0.0414 0.03 0.973

b3 0.0334 0.0414 0.81 0.431

b12 0.0769 0.0498 1.54 0.141

b13 -0.0297 0.0498 -0.60 0.559

b23 -0.0558 0.0505 -1.11 0.284

b11 -0.2269 0.0781 -2.91 0.010

b22 -0.1300 0.0714 -1.82 0.086

b33 -0.0517 0.0714 -0.72 0.479


125

where X is a column vector that contains the nor-

malized values of HNO3 concentration, temperature of

passivation solution and passivation time. Coefficient

of determination R2 for this model reveals that the

passivation parameters explain 81% of variance in

pitting corrosion potential, suggesting that the model

has a good fit. Therefore, the obtained ANN model is

better and more reliable than the MRA model. Details

related to the ANN model are given in Table 3.

Comparison of the models

In order to compare the models as well as effec-

tiveness of the passivation process, the measured

and predicted values of the pitting potential for all

experimental trials are listed in Table 4. Apart from

the results, relative errors of the models were calcul-

ated as well as standard deviation for measured data.

Moreover, the measured pitting potentials for the con-

trol sample are shown in Table 4. Maximal effect of

passivation was measured for passivation condition in

12th experimental trial.

At first, positive effect of the passivation on the

corrosion resistance is obvious according to mea-

sured pitting potential values. Standard deviation of

measured pitting potentials is about of 5% indicating

high reliability of the measured values. Then, taking

into consideration the coefficient of determination for

both models one can notice significantly higher value

for the ANN model. Additionally, MRE shows better

prediction performance of the ANN model since it pro-

duces two times less MRE than the MRA model. In

other words, ANN model is more suitable for the ana-

lysis process with a large non-linearity such as SS

passivation. Therefore, the influence of the passiv-

ation parameters on the corrosion resistance of the

SS is considered by using the ANN model only.

Effects of passivation parameters on pitting potential

The first part of the analysis is concerned with

the analysis of main effects of passivation parameters

on pitting potential. To this aim, Eq. (4) was plotted by

changing one passivation parameter at a time, while

keeping the other two constant at the center level

(Figure 1).

It is evident that the mathematical relationships,

presented graphically in Figure 1, are highly non-

linear. Quantitatively, based on the analysis of the

main effects, concentration is the most influential

parameter, followed by temperature and passivation

time as less influential, respectively. While passivat-

ion temperature and time are on central level, the

highest corrosion resistance is achieved when the

concentration is about 20%. For central levels of con-

centration and passivation time the highest corrosion

resistance is achieved when the temperature is

slightly lower than 30 C. Finally, when the concen-

tration and temperature are set on the central level,

the highest corrosion resistance is achieved when the

duration of the process is about 25 min.

In order to determine the interaction effects of

the passivation parameters on the pitting potential,

3-D surface plots were generated considering two

parameters at a time, while the third one was kept

constant at the center level (Figure 2).

From Figure 2 it can be observed that the pitting

potential is highly sensitive to the selected passiv-

ation parameters. It is also obvious that the effects of

the parameter are variable depending on their own

level, since there are significant interaction effects of

passivation parameters on the pitting potential. The

functional dependence between the pitting potential

and the passivation parameters is strongly nonlinear,

therefore the effect of a given parameter on the pitting

potential must be considered through the interaction

with the other parameters.

For instance, if the passivation time is set on the

central level (40 min), Figure 2a, and temperature on

the low level, an increase in concentration of the sol-

ution leads firstly to the pitting potential increase up to

some extreme value, which corresponds with the

middle level of the concentration. Further increase in

concentration leads to the reduction of the pitting pot-

ential. In this case, the pitting potential is very low

when the concentration is on the high level, while the

middle one provides a fairly high pitting potential.

When the temperature is set on the high level, with an

increase in concentration from the low level the pitting

potential firstly decreases, then increases, and the

nearby high level starts to impair.

When the temperature is set on the middle level

(Figure 2b) and concentration is on the low level, an

Таble 2. ANOVA results for the MRA model;DF - degree of freedom; SS - sum of squares; MS - mean square; F - value of Fisher’s distri-

bution; P - probability density

Source DF SS MS F P

Regression 9 0.60322 0.06702 2.19 0.078

Residual Error 17 0.51944 0.03056

Total 26 1.12267


126

Table 3. The weights and biases of the developed ANN mode; W1: weights between input and hidden layer; W2: weights between hidden

and output layer; B1: biases of the hidden neurons; B2: bias of the output neuron

W1 W2 B1 B2

-0.26021 -2.5744 -0.78867 0.36462 2.1845 -1.2035

-1.2355 -2.2487 -1.317 -1.1196 1.3122

1.4764 1.5512 0.95885 -1.1753 -0.27368

-1.9939 0.97443 0.30408 -1.0401 -2.539

-1 -1.6398 -1.4029 -1.0435 -2.4896

Table 4. Comparative review of the measured pitting potential and predicted values for the pitting potential by means of MRA and ANN

models; important remark: shaded rows - testing data for ANN model performance

Exp. trial Passivation parameters Experimental MRA Model ANN Model

HNO3 concentration

%

Temperature

C

Passivation time

min

Ep

V

Standard deviation

%

Ep

V

Relative error

%

Ep

V

Relative error

%

Control - - - 0.68 4.04 - - - -

1 10 17 20 0.85 4.51 1.08 26.89 0.86 1.15

2 10 17 40 1.16 5.13 1.13 2.21 1.13 2.49

3 10 17 60 1.19 6.11 1.19 0.01 1.47 23.85

4 10 40 20 1.12 4.16 1.19 6.03 1.14 1.37

5 10 40 40 1.44 5.51 1.19 17.53 1.35 6.14

6 10 40 60 1.18 4.58 1.19 0.64 1.22 3.38

7 10 60 20 1.08 3.06 1.04 4.03 1.05 2.49

8 10 60 40 0.77 2.52 0.98 27.35 0.87 12.79

9 10 60 60 0.81 5.20 0.92 14.17 0.78 3.48

10 30 17 20 1.0 4.93 1.29 29.49 1.03 3.00

11 30 17 40 1.39 4.04 1.35 2.83 1.46 5.08

12 30 17 60 1.49 5.03 1.41 5.60 1.47 1.32

13 30 40 20 1.41 3.21 1.46 3.53 1.46 3.27

14 30 40 40 1.47 5.14 1.46 0.70 1.33 9.56

15 30 40 60 1.39 3.04 1.46 5.02 1.30 6.16

16 30 60 20 1.16 5.77 1.36 17.63 1.20 3.37

17 30 60 40 1.41 2.52 1.31 7.18 1.30 7.79

18 30 60 60 1.33 4.62 1.25 5.79 1.26 5.02

19 65 17 20 1.28 4.20 1.10 14.37 1.32 3.35

20 65 17 40 1.29 5.00 1.15 10.71 1.31 1.26

21 65 17 60 0.94 3.55 1.21 28.47 1.04 10.44

22 65 40 20 1.13 2.89 1.36 20.24 1.11 1.94

23 65 40 40 1.14 5.77 1.36 19.18 1.19 4.20

24 65 40 60 1.43 4.93 1.36 4.99 1.48 3.64

25 65 60 20 1.42 3.79 1.36 4.13 1.34 5.40

26 65 60 40 1.17 3.21 1.31 11.59 1.34 14.89

27 65 60 60 1.34 4.15 1.25 6.73 1.31 2.38

Mean relative error (MRE) 11.00 5.53

increase in passivation time leads firstly to the pitting

potential increase up to some extreme value and then

decrease. In this case, the pitting potential is very low

when the passivation time is on the low and high

level, while the middle one provides pretty high pitting

potential. When the concentration is set on the high

level, with increasing passivation time from the low

level the pitting potential firstly increases rashly, then

slightly decreases up to about 30 min, and then starts

to grow again up to the high level.

Based on Figure 2a and b, it can be concluded

that the highest pitting potentials correspond with a

combination of parameters concentration-temperature

and concentration-passivation time slightly below the


127

Figure 1. Main effects of passivation parameters on pitting potential.

Figure 2. Interaction effects of passivation parameters on pitting potential.


128

middle levels. As it can be seen in Figure 2c (where

concentration is on the middle level), the lowest pit-

ting potential predicted for temperature of 17 C and

passivation time 20 min; the highest pitting potentials

is predicted for the temperature and passivation time

slightly below the middle level.

CONCLUSION

Passivation by immersing in nitric acid solution

is an effective method to improve corrosion resistance

of biomedical SS. In biomedical SS passivation pro-

cess, MRA and ANN were introduced to model func-

tional relationship between pitting potential and pas-

sivation parameters such as nitric acid concentration,

temperature and passivation time. A non-linear func-

tional dependence between the passivation para-

meters and the pitting potential was determined.

Hence, the ANN model proved to be more suitable for

modeling the processes such as chemical passivation

of SS. Nitric acid concentration has maximum inf-

luence on the pitting potential followed by the tempe-

rature and passivation time. The best experimental

result was achieved by a combination of parameters:

HNO3 concentration – 30%; temperature – 17 C; pas-

sivation time – 60 min.

Acknowledgement

This paper is a result of the projects ON174004

and TR35034 supported by the Ministry of Education,

Science and Technological Development of the Rep-

ublic of Serbia.

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129

DUŠAN LJ. PETKOVIĆ

MILOŠ J. MADIĆ

GORAN M. RADENKOVIĆ

Univerzitet u Nišu, Mašinski fakultet,

Katedra za proizvodno-informacione

tehnologije, A. Medvedeva 14, 18000

Niš, Srbija

NAUČNI RAD

UTICAJ PARAMETARA PASIVIZACIJE NA PITING POTENCIJAL BIOMEDICINSKOG NERĐAJUĆEG ČELIKA

Pasivizacija predstavlja hemijski process u kome se površina metalnih legura dovodi u

stanje elektrohemijske pasivnosti. Biomedicinski nerđajući čelik AISI 316LVM može se

pasivizirati potapanjem u azotnu kiselinu, jer se time poboljšava zaštitini oksidni sloj;na taj

način se povećava i koroziona postojanost ovog materijala u fiziološkim rastvorima. U

ovom istraživanju, za matematičko modeliranje procesa pasivizacije korišćene su više-

struka regresiona analiza i veštačke neuronske mreže. Za izlazni (zavisni) parametar

modela izabran je piting potencijal, koji predstavlja meru korozione postojanosti. Kao para-

metri pasivizacije razmatrani su: koncentracija azotne kiseline, temperatura i vreme pasi-

vizacije. Upoređeni su eksperimentalni rezultati i rezultati modela. Pokazalo se, da jedino

model dobijen pomoću veštačkih neuronskih mreža ima statistički zadovoljavajuću tačnost

predikcije. Na kraju, na osnovu modela dobijenog pomoću veštačkih neuronskih mreža,

izvedena je analiza uticaja parametara pasivizacije na piting potencijal biomedicinskog

nerđajućeg čelika.

Ključne reči: nerđajući čelik, azotna kiselina, pasivizacija, višestruka regresiona

analiza, veštačke neuronske mreže.






131

MARIJA KODRIC

SANDRA STOJANOVIC

BRANKA MARKOVIC

DRAGAN DJORDJEVIC

University of Niš, Faculty of

Technology, Leskovac, Serbia

SCIENTIFIC PAPER

UDC 677.494.674.027.4:66.081.3:544


MODELLING OF POLYESTER FABRIC DYEING IN THE PRESENCE OF ULTRASONIC WAVES

Article Highlights

• Non-standard dyeing of PES fabric without carrier is possible by ultrasound

• Continual growth of the amount of exhaustion dye with mass of material is noted

• The Langmuir model shows precise description of experimental data

• Kinetic model of pseudo second-order adequately describes disperse dye – PES fab-

ric system

Abstract

In this paper, modelling of dyeing, i.e., adsorptive behaviour of disperse dyes on

polyester fibres under the influence of ultrasound has been considered with the

aim of getting data about binding mechanisms, as well as defining the conditions

of dyeing with additional energy source without the use of carriers, compounds

that increase permeability of the fibres and help dyeing. Dyeing adsorption was

conducted under different conditions, and the concentration of dyes, mass of the

substrate, recipes and time of dyeing were varied. It was established that ultra-

sound allows dyeing without carriers, and that the efficiency of dyeing depends

on the time of contact, initial concentration of the dye and the amount of abs-

orbent material. Continual growth of the amount of bound dye with the mass of

the absorbent was observed. Characteristic plots obtained from confirmed that

the Langmuir isotherm model ensures a precise description of polyester dyeing

by disperse dye. The dyeing kinetics was remarkably well described by pseudo

second-order in regards to the high functionality.

Keywords: adsorption, polyester, disperse dye, ultrasound, Langmuir isotherm, kinetics.

It is known that polyester (PES) belongs to a

group of synthetic fibres that possess active spots

where dye molecules can be adsorbed. PES contains

a great number of ester groups, as well as a certain

number of carboxylic groups, which are placed at the

ends of chains, such that during the dyeing of this

fibre hydrogen bonds with dye molecules will be est-

ablished. PES has significant hydrophobic character

and compact structure, so taking into consideration

this kind of fibres behaviour in the dyeing solution, it is

necessary to modify the usual dyeing process, i.e. to

increase the rates of dye diffusion in the fibres. Gen-

erally, the rates of dyeing could be increased by using

Correspondence: D. Djordjevic, University of Niš, Faculty of

Technology, Bulevar oslobodjenja 124, Leskovac, Serbia. E-mail: [email protected] Paper received: 13 November, 2015 Paper revised: 12 April, 2016 Paper accepted: 13 April, 2016

suitable dyes, changing the fibre structure, changing

the conditions of dyeing, etc. 1,2.

The rate of dye diffusion can be intensified by

increasing the permeability of the fibres, i.e., inc-

reasing the fibre swelling ability. This can be achieved

by adding simple organic compounds (with smaller

molecules than the dye molecules and with certain

affinity towards the fibre) in the dyeing solution. These

compounds are termed as carriers: due to their small

molecular size they quickly diffuse into the fibre, bind

to carboxylic groups and establish hydrogen bridges.

Carriers cannot be completely removed from the fib-

res by washing; therefore, they require special atten-

tion due to their toxic and dermatologic effects. On

the other hand, the presence of these substances on

the fibres adversely affects the fastness of many dyes

with light, and in certain cases they influence fibre

shrinking as well [3,4].

M. KODRIC et al.: MODELLING OF POLYESTER FABRIC DYEING… Chem. Ind. Chem. Eng. Q. 23 (1) 131139 (2017)

132

Ultrasound in the 20-100 kHz frequency range is

used to increase the chemical reaction rate and adv-

ance different physical processes, such as cleaning,

emulsion, extraction, etc. It can achieve the same or

even better results in comparison to already existing

techniques under less extreme conditions (e.g., lower

temperatures and high chemical concentrations of

reactants). Hence, the process of material dyeing

using ultrasound is very significant. Observed imp-

rovements in ultrasound processes of dyeing gen-

erally refer to the phenomenon of cavitation, but some

other mechanical influences can occur too, such as

dispersion and diffusion [5-7].

This paper strives to explain the ability of dye

adsorption into the fibres in aqueous environment

with usual auxiliaries according to standard recipes in

the presence of ultrasound waves, instead of carriers.

The purpose is to successfully perform the process of

dyeing of very hydrophobic and crystal fibres under

normal pressure and temperature conditions. Conse-

quently, if the dye exhaustion is higher, the amount of

dyed wastewater is lower and less harmful to the

environment. The inclination was to eliminate the

usual supplement – carriers, since it is known that

most of these supplements have negative impact on

human health and environment. Furthermore, the

goal was to explain the polyester dyeing process in

new circumstances using a series of experiments, as

well as modelling the system and kinetic parameters.

EXPERIMENTAL

As an adsorbent, 100% polyester fabric (poly-

ethylene terephthalate) has been used with the fol-

lowing characteristics: weave – twill ½S, warp and

weft fineness of 162 tex each, warp and weft density

of 30 and 22 cm-1 and surface mass of 195 g/m2. The

structure of the used disperse dye (adsorbate), C.I.

Disperse Blue 79 (DB79), is shown in Figure 1. It is a

large molecule, highly energetic dye with good fast-

ness in sublimation and wet treatment. It is suitable

for dyeing by exhausting and thermosol procedure, as

well as printing synthetic material.

Figure 1. Structure of the used disperse dye,

C.I. Disperse Blue 79.

The dyeing process (the method of exhaustion)

was performed with the dye bath heated at 60-70 C

with 1.5 g/dm3 of carrier (Icelan PSN, anionic, on the

basis of mixture of different aromatic esters and spe-

cial emulsifiers, Textilcolor AG, Switzerland) and 1

g/dm3 of dispersing agent (TC-Dispergator DTS, non-

-ionic, on the basis of polyglycolether-derivates, Tex-

tilcolor AG, Switzerland). The dye bath was adjusted

to pH 5 by acetic acid, and, after mixing, the disper-

sant substance was added, with continual heating to

95 C. The PES fabric sample was then added and

dyed for 60 min. After dyeing completion, the dyed

material was washed at 70-80 C.

Non-standard dyeing was performed under iden-

tical conditions, but in the absence of carrier. Instead,

an Elac Ultrasonic Laboratory Reactor URS 1000 was

used. The frequency of the applied ultrasound oscil-

lations was 140 kHz, while the power was 50 W.

The amount of PES fabric in the dyeing adsorp-

tion test was varied from 2 to 6 g, and the used dye

solution (constant volume of 0.1 dm3) were prepared

at concentrations of 50, 100, 200, 300 and 400

mg/dm3 in distilled water. The time of treatment, with

constant mixing, was 10, 20, 30, 40, 50 and 60 min.

The time of 60 min was taken as equilibrium no sig-

nificant changes in dyeing adsorption were observed

with further treatment.

For determining the concentration of dye in the

solution, a Cary 100 Conc UV-Vis Varian spectro-

photometer was used (absorption maximum at 550

nm).

The level of dye exhaustion (%) is calculated as

[8]:

0

0

100c c

c (1)

where c0 and c, mg/dm3, are the initial and final con-

centration of the solution of dye.

Langmuir isotherm was obtained from the fol-

lowing equation [9]:

ee

e max max

1 1cc

q Q b Q (2)

where ce, g/dm3, is the equilibrium concentration of

dye after finished adsorption; qe, g/kg, is the adsorbed

amount of adsorbate (dye) per mass unit of the

adsorbent (fabric); Qmax, g/kg, is the maximum

amount of adsorbate which can be bound on the

adsorbent; b, dm3/g, is the Langmuir constant.

The amount of adsorbed dye per mass unit of

fabric is calculated as [9]:


133

0 e

e

c c Vq

w (3)

where w, kg, is the mass of fabric and V, dm3, is the

volume of the solution from which the adsorption was

done.

Lagergren’s pseudo first-order equation is usu-

ally expressed in the linear form [10]:

1e elog log

2.303t

kq q q t (4)

where qe and qt, mg/g, are the capacities of the ads-

orption in equilibrium and after time t, respectively,

and k1, 1/min, is the rate constant of pseudo first-

-order adsorption.

The kinetic equation of the pseudo second-order

adsorption has the form [10]:

2

e2 e

1 1

t

tt

q qk q (5)

where k2, g/mgmin, is the rate constant of the sec-

ond-order adsorption.


The chosen dye DB79 is not water-soluble. The

solubility depends on the chemical composition and

especially on the content of polar and non-polar func-

tional groups. As a typical donor-acceptor chromo-

gen, this dye has two nitro groups and one bromine

that pull on electrons, and, to a certain extent, elec-

tron-donor ethoxy group, whereas acetylamino and

diacetyloxyetilamino groups practically are not suit-

able donors and acceptors of electrons. In consider-

ation of interaction of the mentioned functional groups,

substantivity will occur towards the textile, which is in

this case a hydrophobic material.

The inclusion of dispersing agent in the dye bath

is a crucial factor in the application of disperse dyes.

The hydrophobic tails of the dispersing agent mole-

cules are inside a micelle which, as a consequence,

is able to solubilise the disperse dye molecules, thus

conferring a higher apparent solubility to the dye [11].

The dye transfers to the fibre from the micelles.

As micelles release their dye, they reform and dis-

solve more dye from the solid particles. Much of the

evidence that is available on the subject suggests that

in dyed polyester fibres the disperse dyes are present

chiefly in the monomolecular state. At the end of the

dyeing process, the dye that has been absorbed by

the fibre is in a state of dynamic equilibrium with the

dye that remains in the bath, and the fraction of the

latter that is in aqueous solution must be present in the

same state of aggregation as the dye in the fibre [11].

The influence of time or the length of contact

between adsorbate and adsorbent on adsorption –

dye exhaustion during ultrasound (without carrier), for

different initial concentration of disperse dye, is shown

by comparative plots in Figure 2. The continuity in

changes during time is present, i.e., the longer time

carries greater amount of adsorbed dye per mass unit

of adsorbent. At lower dye concentrations there is

greater dye exhaustion in contrast to higher dye con-

centrations. Realistically, however, higher amount of

dye adsorbed occurs at higher initial dye concentra-

tions.

In other words, for example, dye exhaustion is

72.48% for the concentration of 50 mg/dm3 and 10

min of adsorption, while dye exhaustion is 57.03% for

the concentration of 400 mg/dm3 and 10 min of ads-

orption.

Therefore, there is a greater exhaustion at low

concentration, but after calculating, real dye adsorp-

tion at a lower concentration is: 72.48%50 mg/dm3 =

Figure 2. The influence of adsorption time on the percentage of DB79 exhaustion (A - PES, 2 g; B - PES, 6 g).


134

= 36.24 mg of dye adsorbed on PES fabric; while at

higher concentration the amount is: 57.03%400

mg/dm3 = 228.12 mg of dye adsorbed on PES fabric;

Linear parts of the curve reflect diffusion in the

surface layer, whereas the parts of the plateau on the

curve respond to diffusion in pores. The diagram for

adsorption on 4 g of PES is not shown, because there

are no great differences in the appearance of the

curves.

A similar investigation by Olenka et al. used the

same dye (Disperse Blue 79) for PES dyeing [12].

They implemented the chemical modification of the

substrate, instead of using ultrasound, and PES dye-

ing was performed without carriers. The results of dye

exhaustion showed that the best dyeing conditions

were to treat the PES for 15 min at 85 C with N,N-

-dimethylacrylamide as a modifier, followed by a dye-

ing time of 30 min at 85 C. These conditions were

shown to be suitable for the dye. Photoacoustic spec-

troscopy allows the determination of the penetration

conditions at which the modified PES can absorb

more dye of original.

Also, disperse dye exhaustion was analyzed by

Choi and Kang [13]. They prepared six nano-disperse

dyes using corresponding O/W nanoemulsions in

order to dye PES (two type, regular- and micro-fib-

ers). Dye exhaustion using nano dyes resulted in low

exhaustion yields of 17-26% on regular polyester fiber

and 28-38% on ultramicrofiber polyester. The obs-

erved low exhaustion yields of nano-disperse dye can

be explained by the solubilization of dye particles into

surfactant micelles as well as the high stability of the

nanoemulsions, which might reduce the capacity of

dye uptake by the fibers. As commercial disperse

dyes exhibit exhaustion of 90-95%, these results were

extremely lower than conventional disperse dyes.

However, in the case of dyeing with nano-dyes pre-

pared on ultramicrofibers, it was observed that micro-

fiber site exhibited higher K/S values than those of

regular polyester site, in the range of 1.5-2.4%, which

was promising a possibility for higher efficiency on

ultramicrofibers.

Likewise, the potential of ultrasound as a means

of enhancing dyeing efficiency was evaluated by Lee

and Kim [14]. Changes in the particle sizes of a dis-

perse dye dispersion with ultrasound irradiation are

studied using the turbidity concept, and the effect of

particle size on the exhaustion rate is also inves-

tigated. Ultrasound irradiation of a dye dispersion red-

uces the particle sizes of disperse dyes, and the

exhaustion rate of the dyes on fibers is enhanced by

this reduced particle size by the ultrasound pretreat-

ment before dyeing experiments. These results sug-

gest that ultrasound is useful method of accelerating

the dyeing rate and increasing dyeing efficiency.

During the dyeing of Disperse Blue 56 with and

without ultrasound, as might be expected, increasing

the dyeing temperature increased the dye uptake and

dyeing rate regardless of ultrasound use [15]. Com-

paring the dyeing kinetics of Disperse Blue 56 with

and without ultrasound, authors did not see a signific-

ant increase in dye uptake and dyeing rate for Dis-

perse Blue 56 with ultrasound use. The effect of ultra-

sound on the dyeing behavior of Disperse Blue 56 on

PET fibers was almost non-existent, with little inf-

luence on dye absorption and dyeing rate. PET dye-

ing of Disperse Red 60 with ultrasound irradiation

showed a considerable increase not only in the dye-

ing rate but also in dye uptake. Comparing dyeing

with and without ultrasound, dye uptake was higher

by some 68 and 6%, respectively, for Disperse Red

60 when dyeing was conducted with ultrasound.

Compared with the results for Disperse Blue 56, this

demonstrates that ultrasonic energy can have a pro-

found effect on the dye uptake and dyeing rate of

more crystalline dyes like Disperse Red 60.

A comparison of dyeing with carrier with or

without ultrasound, and dyeing with ultrasound only,

are shown in Figure 3 with respect to their depen-

dency on exhaustion dye–time of dyeing. Only data for

dye concentration of 400 mg/dm3 and minimal and

maximal amount of PES fabric is shown, since very

similar behaviour is observed under other concentra-

tions and 4 g of remaining material mass. The shape

of the curves is very much alike for the given concen-

tration and 6 g of PES, whereas for minimal amount

of PES (2 g), the curves partly differ.

Figure 3. The influence of adsorption-dyeing time on DB79

exhaustion (400 mg/dm3) in different conditions.


135

The fact that stands out is that dyeing with ultra-

sound leads to more exhaustion and binding dye for

PES fibres in comparison to dyeing with carrier with-

out ultrasound. It is interesting that the presence of

ultrasound in the dye bath with all auxiliaries (includ-

ing carrier) gives better results, i.e. the biggest DB79

exhaustion on the fibre. This kind of behaviour prac-

tically is mapped at all dye concentrations and all

amounts of adsorbents. As the amount of adsorbent

is increased, dyeing with ultrasound without carrier

performs much better in comparison to dyeing with

carrier. Nonstandard dyeing with ultrasound, but with

carrier, gives the best results of dyeing in all cases. It

is associated with the fact that the present carrier,

having achieved the ability of swelling, opens the

structure of fibres and causes greater dye exhaustion

that is helped very much by the present ultrasound.

This proves the positive effect of ultrasound

waves on dyeing with and without carrier, which is

explained by dispersion (separation of big dye par-

ticles into smaller ones), then by degassing (ejection

of soluble or trapped air from the capillaries of the

fibres), expansion (speeding up the dye molecules to

penetrate into PES fibres) and intensive mixing of the

dye solution [16].

Aside from the above mentioned, greater dye

exhaustion can be observed due to the change in

PES fibre crystallinity of PES fibres changes during

the ultrasound processing, because of removal of oli-

gomer from the surface and surface layers, and the

cavitation of ultrasound dislocating macromolecules

among micro crystallites, thus enlarging the amor-

phous area [5].

In other words, in order to diffuse the molecules

of dye properly into the fibres, the free volume must

be formed inside the substrate (fibre). The applied

ultrasound helps the free volume to be created easily

inside the polymer via thermal moving of the chains of

molecules and dye molecules, which enter this free

area. At the same temperature, thermal moving of

molecule chains is directly related to the strength of

polymer substrate, i.e., the faster dye diffusion can be

achieved in softer and more flexible substrates of

polymers, which is partially enabled by using ultra-

sound waves [5].

Similarly, Saligram et al. found that sonication

during dyeing brought much more [17]. Namely, ultra-

sound used in the dyeing of polyester fibres with

disperse dyes at low temperature, and results com-

pared with those achievable in conventional dyeing at

the boil using a carrier. Dyeing was enhanced in pre-

sence of carrier and by pre-swelling the fibres,

although the results obtained were not generally as

good as those that can be obtained in conventional

high-temperature processes. Taking into account the

energy conservation aspect, ultrasound appears to be

a promising technique for dyeing.

Also, poly(trimethylene terephthalate) (PTT) fab-

ric, a new type of polyester fibre, was dyed with Dis-

perse Red FB by using ultrasonic power by Wang et

al. [18]. It was shown that the ultrasonic-assisted

dyeing could increase the depth of shade in PTT fab-

ric at a lower temperature. The ultrasonic energy can

disintegrate the large particles of oligomer on the sur-

face of the PTT fibre to smaller ones and slightly dec-

reases the crystallinity of the PTT fibre, which can

reduce the particle size of the disperse dye in the dye

solution as well. Moreover, the ultrasonic dyeing of

the PTT fabric with a swelling agent can enhance the

colour strength of the dyed fabric, thereby reducing

the dyeing time as well as saving energy. The effects

of ultrasound on the K/S values of the fabric, the fibre

structure and the disperse dye were investigated. The

results show that the ultrasonic power increases the

K/S values of the fabric, disintegrates the large par-

ticles of oligomer on the surface of poly(trimethylene

terephthalate) fibre into smaller ones and reduces the

particle size of the dye in solution. The K/S value of

poly(trimethylene terephthalate) fabric dyed using

ultrasound is much higher than that without ultra-

sound, especially at temperatures higher than 60 C.

In a similar manner, an attempt was made to

evaluate the possibility of using ultrasonic techniques

for effective low-temperature dyeing of polyester [19].

The ultrasonic dyeing depends strongly on a preswel-

ling process that would be both expensive and difficult

to carry out commercially (particularly in terms of

health and safety aspects). Performance also dep-

ends on the energy levels of the dyes used. The res-

ults obtained for the dyes of higher relative molecular

mass were much worse than those obtained when

using carrier at the boil (although raising the dyeing

temperature would be expected to provide notable

improvements in depth). Little advantage would be

gained over conventional dyeing methods, particularly

when carrier was incorporated in the dyebath. An

ultrasound dyeing unit with a lower frequency level

(around 26 kHz), to generate more pronounced cavit-

ation effects, may have given better results. However,

the unit chosen had a frequency rating that was both

readily available commercially (38 kHz) and could still

be considered operationally viable. The noise levels

associated with lower frequency units would be unac-

ceptable in commercial use.

Since it is being established that the dye

exhaustion on the polyester is sufficiently good with


136

ultrasound without carrier, the results of modelling of

dyeing – adsorption without carrier but under the inf-

luence of ultrasound are given as follows. The results

of changing the adsorbed amount of adsorbate on the

adsorbent during time, for different initial concentra-

tions of dye during dyeing with ultrasound (without

carrier) in relation to the mass of PES fabric, are

shown in Figure 4. The continuity in changes with

time can be observed, i.e., the longer the time, the

greater the amount of the adsorbed dye per mass unit

of adsorbent. The highest adsorption occurred at the

highest applied concentrations of dye.

Since the total surface area of fibres is higher

than the outside surface, the molecules of dye will

faster adsorb during dyeing than the present auxil-

iaries. Since the dynamic equilibrium of the solution

deranges because of that, the aggregates of dye will

dissolve into molecules and establish equilibrium

again in the solution. Adsorption will continue up to

the point when the equilibrium between dye concen-

tration in the solution and dye concentration in the

fibre is not reached. Since the molecules of dye have

the tendency to form aggregates in aqueous solution,

ultrasound energy causes degradation of dye aggre-

gates in the solution, decreasing the size of dye par-

ticles in dispersion, which is the first pre-condition for

better adsorption on the adsorbent [5-7].

Figure 5 represents the interpretation of Lang-

muir adsorptive isotherm for different amounts of

adsorbents, showing the dependency of parameters

(ce/qe) on the equilibrium dye concentration (ce) during

dyeing with ultrasound (without carriers). From the

slope and intercept of the fits, the values of Langmuir

constants (Qmax and b) were obtained, which are rel-

ated to the maximum amount of dye that can bind on

the fibre, and the free energy of adsorption, respect-

ively.

Figure 5. Langmuir adsorption isotherms for DB79 – PES

fabric system.

Figure 5 shows that adsorption curves are flat

and continual, which leads to saturation during dif-

ferent concentrations on the outer interphase of ads-

orbent material. Moderate dispersing of data is pre-

sent which indicates the adequacy of Langmuir iso-

therm for describing the adsorption equilibrium of the

examined systems. This kind of behaviour could be

explained by the assumption that the dye in the

beginning adsorbs on the outer surface of PES fabric,

and after reaching a certain level of saturation, it

enters the inner space of PES fibres, when it becomes

adsorbed by inner surfaces. After diffusing of dye into

pores of the fibres, diffused resistance increases,

which leads to a decrease in diffusion rate. By dec-

reasing the concentration of dye in the solution, the

diffusion rate becomes constantly lower so the dif-

fusion process reaches an equilibrium state, which

completely submits to the law of equilibrium adsorpt-

ion defined by Langmuir isotherm.

Figure 4. Adsorbed amount of dye DB79 on PES fabric during time of ultrasonic treatment (A - PES, 2 g; B - PES, 6 g).


137

Table 1 shows the analytical expressions of

Langmuir isotherm, values of Langmuir parameters

Qmax and b, and values of coefficient of determination

R2. The coefficient of determination is a relative mea-

sure of representability of regression line or measures

of usage of Langmuir model.

According to the results for coefficient determin-

ation from Table 1, high values (above 0.97) are

noted, which means that great percentage of square

sums of variables values that deviate from arithmetic

mean is explained by regression model. Therefore, in

the procedure of determining capacities of dye retain-

ing and affinities of adsorbent to dye, Qmax and b,

respectively, the Langmuir monolayer model fits the

experimental data reasonably well and it can be

acceptable for adsorption of disperse dye on PES

fabric.

The Langmuir model parameters are strongly

dependent on the amount of adsorbent: Qmax dec-

reases with the increasing of PES amount, whereas

the values of other constant, b, increase continually

with increasing the mass of the material. Higher

values of b obtained for PES – disperse dye system

mean stronger binding of dye.

In the study of Wang et al., the Nernst, Langmuir

and Freundlich isotherm models were employed in

the fitting of the experimental points using the soft-

ware Origin [18]. It can be seen that the coefficients of

determination (R2) of the Langmuir isotherm are the

highest in the three models, indicating that the Lang-

muir model is the most effective in simulating the ads-

orption isotherm of Disperse Red FB onto the PTT

fibre. In addition, the adsorption capacities (Qmax) of

the PTT fibre dyed in the presence of ultrasound are

all higher than that of the PTT fibre dyed without ultra-

sound at three temperatures. This proves the ultra-

sound-assisting effect on the dyeability of PTT fibre

with disperse dye.

Figure 6 shows diagrams with results related to

kinetic sorption of DB79 on PES fabric for the applied

mass of 2 g of adsorbent and different initial dye con-

centration during dyeing with ultrasound (without car-

rier). According to linear forms of pseudo first-order

model (Figure 6A), it can be concluded that the rates

of adsorption in given experimental conditions, does

not describe properly pseudo first-order model for the

whole period of sorption, in comparison to the model

of pseudo second-order (Figure 6B) which gives func-

tional straight line for all initial concentrations of dye.

Based on this, it can be said that, in this particular case,

the pseudo second-order model is more usable. Simi-

lar behaviour is observed for mass of material of 4

and 6 g, hence those results are not shown in this text.

Tables 2 and 3 show results of kinetic para-

meters of process of adsorption DB79 on PES fabric

(equilibrium rates constant of kinetics for pseudo first-

-order and pseudo second-order) for the used mass

of adsorbent, all initial concentration of dye, as well as

values for parameter q (calculated – qcal and experi-

mental – qexp).

Table 1. Analytical expressions of Langmuir isotherm with coefficient for DB79 – PES system

Adsorbent, g Analytical expression Langmuir’s parameters R2

(Langmuir’s equations) Qmax / mg g–1 b / dm3 mg–1

2 ce/qe = 9.635 + 0.078ce 12.67 0.008 0.973

4 ce/qe = 12.295 + 0.175ce 5.70 0.014 0.992

6 ce/qe = 9.057 + 0.180ce 5.55 0.020 0.997

Figure 6. Kinetic of sorption DB79 on 2 g of PES fabric (A - pseudo first-order; B - pseudo second-order).


138

Table 2. Kinetic parameters of process of adsorption of dye on

2 g of PES fabric (pseudo first-order)

Dye concentration

mg/dm3

qe,exp

mg/g

Pseudo first-order

k1 / min–1 qe,cal / mg g–1 R2

50 0.98 0.074 0.078 0.921

100 1.93 0.145 0.295 0.979

200 3.83 0.078 0.831 0,949

300 5.29 0.070 0.912 0.933

400 6.60 0.066 0.922 0.965

Table 3. Kinetic parameters of process of adsorption of dye on

2 g of PES fabric (pseudo second-order)

Dye concentration

mg/dm3

qe,exp

mg/g

Pseudo first-order

k2 / g mg–1 min–1 qe,cal / mg g–1 R2

50 0.98 3.63 0.98 0.999

100 1.93 1.63 1.95 0.999

200 3.83 0.32 3.86 0.998

300 5.29 0.28 5.32 0.998

400 6.60 0.28 6.61 0.999

Although the coefficient of determination R2 for

kinetic model of pseudo first-order is higher than 0.92,

completely different values are obtained for the cal-

culated parameter (qcal) in comparison to those for the

experimental parameter (qexp). Because of this, ads-

orption cannot be well described by the kinetic model

of pseudo first-order, because in many cases the

equation of pseudo first-order does not cover ade-

quately the whole range of contact time, which is con-

firmed by the results in Table 2.

In contrast, kinetic model of pseudo second-

order has in all cases R2≈1, Table 3, by which the

complete functionality is achieved and the model can

be completely used for describing processes of ads-

orption of dye on PES fabric. In addition, differences

between parameters qcal and qexp for this model are

insignificant, i.e. acceptable.

Concentration of dye decreases very fast during

initial absorption, before diffusion inside the particles

starts to control adsorption kinetics in all cases. Inc-

rease of contact time decreases the resistance of

borderline layers, and supported by ultrasound waves,

it intensifies the mobility of dye during the time of ads-

orption [20].

In addition, the paper of Carrion-Fité proposes a

dyeing process for polyester at low temperatures (40

C and below) with disperse dyes using a microemul-

sion prepared by ultrasonic agitation, composed of a

small proportion of organic solvent (alkyl halogen)

and phosphogliceride as the emulsifier [21]. The kin-

etics of this dyeing system are determined as a funct-

ion of temperature with various disperse dyes at dif-

ferent molecular weights. In general, dyes with lower

molecular weight have a faster dyeing rate and those

with a higher molecular weight have a lower rate.

Activation energies range from 20-40 kcal/mol; these

values are similar to those achieved in traditional dye-

ing with a carrier.

CONCLUSIONS

Modelling of disperse dye adsorption, i.e., dye-

ing of PES fabric was being tested under different

conditions without carrier in dye bath, although it is

well-known that dyeing of synthetic PES by disperse

dye is performed usually in the presence of this che-

mical and/or in conditions of high pressure and tem-

perature. Usually, the used carriers can show prob-

lems in usage for such dyed material or in production

itself during dyeing. Use of ultrasound could solve all

the drawbacks, having in mind health advantages and

economic savings.

Based on obtained experimental results, the fol-

lowing conclusions can be stated:

• non-standard dyeing of PES fabric without

carrier is possible by ultrasound in atmospheric con-

ditions;

• the longer the contact time, the higher the

amount of dye adsorbed on the material;

• the concentration of dye in solution dec-

reases with the duration of adsorption;

• the continuity of growth in the amount of

exhaustion dye with mass of material is noted, i.e.,

greater adsorption occurs in smaller proportion of bath;

• data obtained from teh Langmuir equation

shows that this model enables precise description of

experimental data;

The kinetic model of pseudo second-order ade-

quately describes disperse dye – PES fabric system.

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Analyst 127 (2002) 310–314

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(1996) 245-248

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MARIJA KODRIĆ

SANDRA STOJANOVIĆ

BRANKA MARKOVIĆ

DRAGAN ĐORĐEVIĆ

Univerzitet u Nišu, Tehnološki fakultet,

Bulevar oslobođenja 124, Leskovac,

Srbija

NAUČNI RAD

MODELOVANJE BOJENJA POLIESTARSKE TKANINE U PRISUSTVU ULTRAZVUČNIH TALASA

U radu je razmotreno modelovanje bojenja, odnosno, adsorpciono ponašanje disperzne

boje na poliestru (bojenje) pri delovanju ultrazvuka, sa ciljem dobijanja podataka o meha-

nizmu vezivanja boje i definisanju uslova procesa bojenja ovog sintetičkog vlakna uz

dodatni izvor energije bez primene kerijera, jedinjenja koja povećavaju permeabilnost vla-

kana i pomažu bojenje. Bojenje-adsorpcija je vođena pod različitim uslovima, varirana je

koncentracija boje, masa supstrata, receptura i vreme bojenja. Utvrđeno je da ultrazvuk

dozvoljava bojenje bez kerijera a efikasnost bojenja zavisi od vremena kontakta, početne

koncentracije boje i količine adsorbenta - tkanine. Postoji kontinuitet rasta količine vezane

boje sa masom adsorbenta. Karakteristični prikazi dobijeni iz Langmuir-ove izoterme potvr-

dili su da ovaj model obezbeđuje dovoljno precizan opis bojenja poliestra disperznom

bojom. Kinetika bojenja odlično je protumačena pseudo II redom s obzirom na visoku funk-

cionalnost.

Ključne reči: adsorpcija, poliestar, disperzna boja, ultrazvuk, Langmuir izoterma,

kinetika.






141

AISHI ZHU

SHANSHAN LIU

KANFENG WU

CHUAN REN

MAOQIAN XU

School of Biological and Chemical

Engineering, Zhejiang University of

Science and Technology,

Hangzhou, China

SCIENTIFIC PAPER

UDC 633.17:577.114.4:66.061:66


COMPARING OF HOT WATER AND ACID EXTRACTION OF POLYSACCHARIDES FROM PROSO MILLET

Article Highlights

• Polysaccharides were extracted from proso millet with hot water and acid solution

• Response surface methodology was used and a model was set up

• The best extraction methods were obtained

• The polysaccharides yield of acid extraction was significantly higher than hot water

extraction

Abstract

The extraction of polysaccharides from proso millet was investigated experi-

mentally using hot water and acid aqueous solution. Response surface method-

ology, based on a three-level, three- or four-variable Box-Behnken design for hot

water extraction or acid extraction, respectively, was employed to obtain the best

possible combination of acid concentration, liquid-solid ratio, extraction time, and

extraction temperature for maximum polysaccharides yield. The obtained experi-

mental data were fitted to a second-order polynomial equation and analyzed by

appropriate statistical methods. The corresponding optimum extraction condi-

tions of each method were obtained. Under the optimum conditions, the experi-

mental yield was well in close agreement with the predicted value by the model.

The results showed that the polysaccharides yield of acid extraction was 42.13

mg g-1, significantly higher than 20.07 mg g-1 of the yield of hot water extraction,

the obtained equation could be used to predict the extraction experimental

results.

Keywords: proso millet, polysaccharides, hot water extraction, acid ext-raction, response surface methodology.

The proso millet (Panicum miliaceum L.) is a

warm-season grass with a short growing season and

low moisture requirement that is capable of producing

food or feed where other grain crops would fail. It is

the a widely planted species of millet; is extensively

cultivated for its grain in the arid areas of China, India,

Russia, Ukraine, the Middle East, Turkey and Rom-

ania. The seeds as grain are small (2–3 mm) and rich

in starch, protein, fat, dietary fiber, vitamin and trace

elements. It has many functions beneficial to human

health, such as the prevention and management of

diabetes mellitus [1], callus and shoots regeneration

Correspondence: A. Zhu, School of Biological and Chemical

Engineering, Zhejiang University of Science and Technology,

318 Liuhe Road, Hangzhou 310023, China. E-mail: [email protected] Paper received: 11 October, 2015 Paper revised: 2 January, 2016 Paper accepted: 26 April, 2016

from protoplasts of proso millet [2]. Polysaccharides

have unique biological properties such as anti-oxid-

ative [3], anti-viral [4] and anti-complementary acti-

vities [5], as well as chemical and physical properties

[6]. They have important effects in the process of

growth and development for living organisms [7].

Response surface methodology (RSM) is an

effective statistical technique for optimizing complex

processes. It is widely used in optimizing the extract-

ion process variables [8,9].

Kim et al. [1] studied the inhibitory effects of

ethanol extracts from proso millet on α-glucosidase

and α-amylase activities, Heyser [2] studied the callus

and shoot regeneration from protoplasts of proso

millet, Yañez et al. [10] studied some chemical and

physical properties of proso millet starch. However,

there is no report about polysaccharides extraction

from proso millet, especially extraction with acid sol-

A. ZHU et al.: COMPARING OF HOT WATER AND ACID EXTRACTION… Chem. Ind. Chem. Eng. Q. 23 (1) 141150 (2017)

142

ution. In this article, we report on the extraction of

polysaccharides from proso millet by hot water and

acid solution. Polysaccharides extracted with the

acidic aqueous solution can be made for a pure poly-

saccharide, greatly improve the yield of polysacchar-

ide, and the activity of polysaccharide is higher [11],

high acid concentration can accelerate the degrad-

ation of polysaccharides [8]. Based on the results of

the single factor investigation, once extracted and stir-

ring rate of 150 rpm in the extraction process, RSM

was designed to help to optimize extraction variables

of acid concentration (hydrochloric acid aqueous sol-

ution, mol L-1), liquid-solid ratio (water or acid aque-

ous solution volume with proso millet mass, mL g-1),

extraction time (h) and extraction temperature (C)

and to systematically analyzed the influence of the

variables.


Materials and extraction of polysaccharides

The decorticated grain proso millet was obtained

from Donghua rice industry limited company, Jinzhou,

Liaoning, China, 2014 harvest. The size of proso mil-

let particle was about 2–3 mm. Before the extraction

experiment, the sample was placed first in electro-

thermal blast oven (DHG-9123A, Precision Experi-

mental Facilities Limited Company, Shanghai, China)

to dry at 60.0 C to constant mass. Five grams of dry

proso millet, weighed with an electronic balance

(BS124S Sartorius Instrument System Limited Com-

pany, Beijing, China) and mixed with different volume

of water or different concentration acid solution was

extracted in a 500-ml three-neck flask. The three-

-neck flask was soaked in a thermostatic water bath

(DK-S24 Jing Hong Experimental Facilities Limited

Company, Shanghai, China) which controlled the

needed extraction temperature, with stirring at 150

rpm with an electric mixing paddle for a given time

during the entire extraction process. After extraction,

the mash was quickly put in cold bath, cooled down to

room temperature and then vacuum-filtered. The fil-

trate was concentrated in a rotary evaporator (RE52CS

Yarong Biochemistry Instrument Plant, Shanghai,

China) to 20% of the initial volume at 60.0 C under

vacuum. The obtained solution was mixed with four

volumes of dehydrated ethanol (ethanol final concen-

tration, 80%) to obtain the precipitate. Then, the sus-

pension was centrifuged using centrifuge (800B

Shanghai Anting Scientific Instrument Factory, Shan-

ghai, China) at 4000 rpm for 15 min, the precipitate

was collected as extract and washed three times with

dehydrated ethanol. The extract was dried at 50 C in

electrothermal blast oven until constant mass and

weighed. The concentration of polysaccharides con-

tent in the extract was examined with a spectrophoto-

meter (722E Shanghai Spectrum Instruments Co.,

LTD, Shanghai, China). The proso millet was ext-

racted once and the extraction procedure was rep-

eated twice. All reagents were of analytical grade.

Experimental design

The Box-Behnken Design was applied to deter-

mine the best combination of extraction variables for

the yield of proso millet polysaccharides. A three-

level, three-variable experimental design was carried

out to hot water extraction, and three extraction vari-

ables considered for this research were: liquid-solid

ratio (mL g-1, X1), extraction time (h, X2), and extract-

ion temperature (C, X3). A three-level, four-variable

experimental design was used to acid extraction, and

four extraction variables considered for these

researches were: acid concentration (mol L-1, X1),

liquid-solid ratio (mL g-1, X2), extraction time (h, X3),

and extraction temperature (C, X4) [8]. The reason-

able range of variables was obtained through the

single factor investigation. The polysaccharides ext-

raction yield was as the dependent variable. The

complete experiment scheme was consists of 17 or

29 experimental points (including five replicates of the

center point) for three-level, three-variable experi-

mental design or three-level, four-variable experimen-

tal design, respectively. A quadratic polynomial model

was used to fit the obtained experimental data and

the regression coefficients were obtained through sta-

tistical analyses. The mathematic expression of the

quadratic polynomial model was seen Eq. (1) [8,12-

–14]:

4 4 3 4

20

1 1 1 1i i ii i ij i j

i i i j j

Y X X X X (1)

where Y is the measured response (polysaccharides

yield) of each experiment; β0, βi, βii, βij are constant

regression coefficients of the model; X is the code

levels of independent variables; XiXj and Xi2 represent

the interaction and quadratic terms of independent

variables, respectively.

Analysis of samples

The phenol-sulphuric acid colorimetric method

at 485 nm was used to determine the polysaccharides

content [12,13]. In this method the glucose was used

as standard, the polysaccharides yield (mg g-1) was

calculated using Eq. (2):

1000XV

YM

(2)


143

where Y is yield of polysaccharides, mg g-1; V is the

total volume of the coarse polysaccharides dissolved

after the constant volume, L; M is dosage of raw

material millet, g; X is concentration of polysacchar-

ides in solution, mg L-1. X was calculated using Eq.

(3):

0.0255 0.00006X A (3)

where A is the absorbance of measured solution at

485 nm.

Statistical analyses

The Design Expert Software (version 8.0.5.0,

Stat-Ease Inc., Minneapolis, MN) was used to mul-

tiple nonlinear regressions of the responses obtained

from each design experimental. The fitting degree of

between the experimental data and the equation was

inspected using the coefficient of determination R2,

F-test and p-value were used for checking the sig-

nificance of the regression coefficient. p-Values below

0.05 were regarded as statistically significant.


Hot water extraction

Model fitting. The factors and their levels of hot

water extraction were chosen on the basis of single-

factor experiments (Table 1). The experiment design

and the yields are shown in Table 1. The regression

analysis results are shown in Table 2.

Table 1. Experimental design and results of hot water extraction

Test number Liquid-solid ratio Extraction time Extraction temperature

Y / mg g-1 X1 / mL g-1 X2 / h X3 / C

1 20:1 1.0 65 7.85±0.22

2 20:1 1.0 75 16.11±0.33

3 20:1 2.0 65 6.20±0.29

4 20:1 2.0 75 17.91±0.33

5 15:1 1.5 65 4.14±0.21

6 15:1 1.5 75 17.19±0.25

7 25:1 1.5 65 7.63±0.27

8 25:1 1.5 75 15.24±0.29

9 15:1 1.0 70 16.52±0.34

10 15:1 2.0 70 14.88±0.31

11 25:1 1.0 70 14.75±0.27

12 25:1 2.0 70 17.94±0.21

13 20:1 1.5 70 19.82±0.22

14 20:1 1.5 70 19.20±0.22

15 20:1 1.5 70 19.46±0.22

16 20:1 1.5 70 19.34±0.22

17 20:1 1.5 70 19.27±0.22

Table 2. Analysis of variance for regression equation to hot water extraction; R2 = 0.9987, Adj. R2 = 0.9970; **extremely significant;

*significant

Source Sum of squares df Mean square F-value Prob > F Significance

Model 423.54 9 47.06 597.64 < 0.0001 **

X1 1.00 1 1.00 12.71 0.0092 **

X2 0.36 1 0.36 4.59 0.0694 –

X3 206.35 1 206.35 2620.53 < 0.0001 **

X1X2 5.83 1 5.83 74.07 < 0.0001 **

X1X3 7.40 1 7.40 93.96 < 0.0001 **

X2X3 2.98 1 2.98 37.79 0.0005 **

X12 20.04 1 20.04 254.47 < 0.0001 **

X22 6.21 1 6.21 78.81 < 0.0001 **

X32 161.15 1 161.15 2046.50 < 0.0001 **


144

Table 2. Continued


Residual 0.55 7 0.079 –

Lack of fit 0.31 3 0.10 1.74 0.2962 –

Pure error 0.24 4 0.060 –

Cor. total 424.09 16 –

The response surface plots graphs are shown in

Figure 1. The response value Y (i.e., the proso millet

polysaccharides yield) can be expressed by Eq. (4) in

terms of actual values:

1 2 3

1 2 1 3 2 3

2 2 21 2 3

1337.566 6.645 18.817 36.231

0.483 0.054 0.345

0.087 4.856 0.247

Y X X X

X X X X X X

X X X

(4)

The results of the analysis of variance, good-

ness-of-fit and the adequacy of the models are sum-

marized in Table 2. The value of probability (p) was

less than 0.05, which indicates that the selected

factors and their ranges have significant influence on

the yield of polysaccharides. The residual analysis

was then performed to check the adequacy of the

developed model and determine whether the approx-

imating model would give poor or misleading results.

Figure 2 shows the residual and the influence plots

for the experimental data [9]. The predicted values

obtained are quite close to the experimental values,

and the points of all predicted and experimental

response values fall very close to the 45 line (Figure

2a), indicating that the model developed is successful

in capturing the correlation between the process vari-

ables on the response. Figure 2b shows the normal

probability plot of residuals for response is normally

distributed, as they lie reasonably close on a straight

line and shows no deviation of the variance. The

goodness of fit of the model is analyzed by construct-

ing the internally studentized residuals versus experi-

mental runs and shows that all the data points lay

within the limits (Figure 2c). Since the Cook’s dis-

tance values are in the determined range (Figure 2d),

there is no strong evidence of influential observations

in experimental data. The above results indicate a

good adequate agreement between BBD experimen-

tal data and the model could be better predicted yield

of polysaccharides. Linear term of X3 (extraction tem-

perature, p < 0.0001) showed the largest effect on

polysaccharide yield, followed by linear term of X1

(liquid-solid ratio, p = 0.0092 < 0.05) , all interaction

terms and all quadratic terms (p < 0.0001) were also

extremely significant. Linear term of X2 (extraction

time, p = 0.0694) was however not significant (p >

> 0.05). The p-value of model was less than 0.0001

and Adj R2 was 0.9970 which would give a better fit to

the mathematical model (Eq. (4)).

Figure 1. a) Response surface plot of Y = f(X1,X2) (X3 = 70.0

C); b) response surface plot of Y = f(X1,X3) (X2 = 1.5 h);

c) Response surface plot of Y = f(X2,X3) (X1 = 20.0 mL g-1);

X1: liquid-solid ratio (mL g-1), X2: extraction time (h), X3:

extraction temperature (C), Y: yield (mg g-1).


145

Verification of the model. The regression model

predicted the optimum extraction conditions of proso

millet polysaccharides, which were liquid-solid ratio

19.77:1 mL g-1, extraction time 1.85 h, extraction tem-

perature 71.79 C. Under this optimum extraction

conditions, the polysaccharides yield was predicted

for 20.17 mg g-1.

The experiment was repeatedly carried out three

times for verifying the prediction from the model at

liquid-solid ratio 19.8:1 mL g-1, extraction time 1.8 h

and extraction temperature 71.8 C. The average

value of practical polysaccharides yield was 20.07±

±0.21 mg g-1(n = 3), the relative error was -0.5%

compared with predicted yield of 20.17 mg g-1. The

analysis results proved that the quadratic polynomial

model was suitable for expressing the optimization

results and the satisfaction and accuracy of Eq. (4)

was high.

Acid extraction

Model fitting. The factors and their levels of acid

extraction were chosen on the basis of single-factor

experiments (Table 3). The experiment design pro-

posal and the yields are shown in Table 3. The reg-

ression analysis results are shown in Table 4.

Table 3. Experimental design and results of acid extraction

Test number Acid concentration Liquid-solid ratio Extraction time Extraction temperature

Y / mg g-1 X1 / mol L-1 X2 / mL g-1 X3 / h X4 / C

1 3.0 20:1 1.5 70 30.48±0.53

2 2.5 25:1 1.5 60 12.08±0.26

3 2.0 25:1 1.5 70 22.67±0.35

4 2.5 25:1 1.0 70 22.94±0.45

Figure 2. Diagnostic plots for the model adequacy of hot water extraction; a) predicted vs. actual, b) normal plot of residuals, c) residuals

vs. run and d) Cook’s distance.


146

Table 3. Continued

Test number Acid concentration Liquid-solid ratio Extraction time Extraction temperature

Y / mg g-1 X1 / mol L-1 X2 / mL g-1 X3 / h X4 / C

5 3.0 20:1 2.0 70 26.86±0.58

6 2.5 20:1 1.0 80 31.00±0.54

7 2.5 15:1 2.0 70 25.64±0.42

8 2.5 15:1 1.5 60 12.88±0.26

9 2.0 20:1 1.5 60 10.24±0.36

10 2.0 20:1 1.0 70 17.88±0.28

11 3.0 20:1 1.5 60 26.87±0.37

12 3.0 20:1 1.5 80 31.88±0.48

13 2.0 20:1 1.5 80 35.50±0.63

14 2.5 25:1 2.0 70 27.50±0.53

15 2.5 25:1 1.5 80 34.99±0.57

16 2.5 20:1 2.0 80 32.09±0.42

17 2.5 15:1 1.0 70 25.58±0.34

18 2.0 20:1 2.0 70 29.97±0.38

19 2.5 20:1 2.0 60 14.98±0.29

20 3.0 25:1 1.5 70 30.48±0.39

21 3.0 15:1 1.5 70 27.27±0.47

22 2.5 15:1 1.5 80 31.61±0.50

23 2.0 15:1 1.5 70 23.55±0.41

24 2.5 20:1 1.0 60 15.39±0.45

25 2.5 20:1 1.5 70 40.67±1.16

26 2.5 20:1 1.5 70 37.59±1.16

27 2.5 20:1 1.5 70 38.25±1.16

28 2.5 20:1 1.5 70 38.52±1.16

29 2.5 20:1 1.5 70 40.07±1.16

Table 4. Analysis of variance for regression equation to acid extraction; R2 = 0.9641, Adj. R2 = 0.9281; **extremely significant;

*significant


Model 2011.80 14 143.70 26.82 <0.0001 **

X1 77.36 1 77.36 14.44 0.0020 **

X2 0.099 1 0.099 0.019 0.8936 –

X3 17.07 1 17.07 3.19 0.0960 –

X4 913.18 1 913.18 170.42 <0.0001 **

X1X2 4.20 1 4.20 0.78 0.3911 –

X1X3 37.06 1 37.06 6.92 0.0198 *

X1X4 102.48 1 102.48 19.12 0.0006 **

X2X3 14.06 1 14.06 2.62 0.1276 –

X2 X4 4.44 1 4.44 0.83 0.3782 –

X3 X4 0.56 1 0.56 0.11 0.7504 –

X12 213.43 1 213.43 39.83 <0.0001 **

X22 316.17 1 316.17 59.00 <0.0001 **

X32 320.45 1 320.45 59.80 <0.0001 **

X42 448.60 1 448.60 83.72 <0.0001 **

Residual 75.02 14 5.36 – – –

Lack of fit 68.32 10 6.83 4.08 0.0938 –

Pure error 6.70 4 1.67 – – –

Cor. total 2086.82 28 – – – –


147

The response surface plots are shown in Figure

3. The response value Y can be expressed by the

following second order polynomial equation in terms

of actual values:

1 2

3 4 1 2

1 3 1 4 2 3

22 4 3 4 1

2 2 22 3 4

906.997 200.730 7.565

96.915 14.512 0.410

12.175 1.012 0.750

0.021 0.075 22.944

0.279 28.115 0.083

Y X X

X X X X

X X X X X X

X X X X X

X X X

(5)

The results of the analysis of variance, good-

ness-of-fit and the adequacy of the models are sum-

marized in Table 4. Figure 4 shows the residual and

the influence plots for the experimental data, it is

similar to Figure 2. The linear term of X1 (acid con-

centration, p = 0.002) and X4 (extraction temperature,

p < 0.0001) had an extremely significant effect on

polysaccharide yield. Each quadratic term (p <

< 0.0001) and the interaction term of X1 and X4 (p =

= 0.0006) were also extremely significant. The inter-

Figure 3. Response surface plots of: a) Y = f(X1,X2) (X3 = 1.5 h, X4 = 70.0 C); b) Y = f(X1,X3) (X2 = 20 mL g-1, X4 = 70.0 C);

c) Y = f(X1,X4) (X2 = 20 mL g-1, X3 = 1.5 h); d) Y = f(X2,X3) (X1 = 2.5 mol L-1, X4 = 70.0 C); e) Y = f(X2,X4) (X1 = 2.5 mol L-1, X3 = 1.5 h);

f) Y = f(X3,X4) (X1 = 2.5 mol L-1, X2 = 20 mL g-1); X1: acid concentration (mol L-1), X2: liquid-solid ratio (mL g-1), X3: extraction time (h),

X4: extraction temperature (C), Y: yield (mg g-1).


148

action term of X1 and X3 (acid concentration and ext-

raction time, p = 0.0198 < 0.05) was significant. The

other terms were however not significant (p > 0.05).

The p-value of model was less than 0.0001 and Adj.

R2 was 0.9281 which would give a better fit to the

mathematical model (Eq. (5)).

Verification of the model. The optimum con-

ditions of acid extraction predicted by regression

model were acid concentration 2.48 mol L-1, liquid-

-solid ratio 20.3:1 mL g-1, extraction time 1.56 h and

extraction temperature 75.45 C. Under these opti-

mum extraction conditions, the polysaccharides yield

was predicted for 41.41 mg g-1.

The experiment was repeated three times at

acid concentration 2.48 mol L-1, liquid-solid ratio

20.3:1 mL g-1, extraction time 1.56 h and extraction

temperature 75.5 C. The average value of practical

yield of polysaccharide was 42.13±0.15 mg g-1(n=3),

the relative error was 1.74% compared with predicted

yield of 41.41 mg g-1. The analysis results confirmed

that the second order polynomial equation was

adequate for reflecting the expected optimization and

the Eq. (5) was satisfactory and accurate.

Effect of acid on yield of polysaccharides

The experimental results showed that the poly-

saccharides yield of acid extraction was significantly

higher than that of hot water extraction. Extraction

yield was increased 109.9%. This is because the acid

can help eliminate the physical and chemical effect

between cell wall of polymer molecules, make more

polysaccharides to dissolve from cells into solution in

the same extraction time, and thus the yield of poly-

saccharides is increased. However, high acid concen-

tration would reduce the polysaccharides yield because

of the destruction of the structure of polysaccharide

caused by the acid catalyzed hydrolysis [8].

Effect of other factors on yield of polysaccharides

The obtained experiment data indicated that

liquid-solid ratio, extraction time and extraction tem-

perature of each extraction method were similar, and

their trends of influence on polysaccharides yield in

Figure 4. Diagnostic plots for the model adequacy of acid extraction; a) predicted vs. actual, b) normal plot of residuals, c) residuals vs.

run and d) Cook’s distance.


149

each extraction process were also similar. The yield

raised with increasing liquid-solid ratio because the

polysaccharides concentration difference between

sample and solution was increased to result in the

raised of mass transfer driving force. However, if the

liquid-solid ratio was too high, it would lead to inc-

reased operating costs, and the increased of amount

of acid in the extract system would lead to the des-

truction of the structure of polysaccharide caused by

the acid catalyzed hydrolysis. These results are in

good agreement with previous observations [8,13-15].

The extraction time had a similar effect on yield as

liquid-solid ratio. The yield first increased and then

decreased with increasing of the extraction time. This

phenomenon could be explained by that the proso

millet cell-wall was broken, the liquid was infiltrated

into the dried sample, the polysaccharides in sample

was dissolved and subsequently diffused out from the

sample to exterior solvent, which needed a long time,

but the polysaccharides dissolved in solution would

be partly degradation because they remained in the

solution for a long time [13,16]. The extraction yield

was found to increase with increase of extraction tem-

perature, and then decreased after a peak. The rea-

son was that the increased system temperature res-

ulted in the decreased of solvent viscosity to enhance

the solvent and solute diffusivity within suspension

solution system, which raised the polysaccharides

solubility. But the polysaccharide could be degraded

under high temperature, so the yield would be dec-

reased [13,15]. There were no obvious differences

between Figures 2 and 4, both shown that no obvious

patterns were found in the analysis of model and

indicated the accuracy of the developed model.

CONCLUSION

The performance of the extraction of polysac-

charides from proso millet was studied with a sta-

tistical method based on the response surface meth-

odology in order to identify and quantify the variables

that may maximize the yield of polysaccharides. The

second order polynomial model had higher correlation

for experiment data and could be better used for

optimizing proso millet polysaccharides extraction

technology. The optimum conditions of hot water ext-

raction were liquid-solid ratio 19.8:1 mL g-1, extraction

time 1.8 h and extraction temperature 71.8 C, and

the yield was 20.07±0.21 mg g-1 (n = 3). The optimum

conditions of acid extraction were acid concentration

2.48 mol L-1, liquid-solid ratio 20.3:1 mL g-1, extraction

time 1.56 h and extraction temperature 75.5 C, the

polysaccharides yield was 42.13±0.15 mg g-1 (n = 3).

The acid solution extraction of polysaccharides was

better than hot water extraction; the yield of acid sol-

ution extraction was higher by 109.9% than of hot

water extraction. The optimal extraction conditions of

hot water extraction and acid extraction were deter-

mined, and under the optimum conditions, the prac-

tical yield was agreed closely with the predicted yield

value.

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150

AISHI ZHU

SHANSHAN LIU

KANFENG WU

CHUAN REN

MAOQIAN XU

School of Biological and Chemical

Engineering, Zhejiang University of

Science and Technology, Hangzhou,

China

NAUČNI RAD

POREĐENJE EKSTRAKCIJA POLOSAHARIDA IZ PROSA TOPLOM VODOM I KISELIM RASTVOROM

Ispitivana je ekstrakcija polisaharida iz prosa pomocu tople vode i kiseli vodeni rastvor. Za

određivanje najbolje moguće kombinacije koncentracije kiseline, odnosa tečno-čvrsto, vre-

mena i temperature ekstrakcije koja obezbeđuje ostvarivanje najvećeg prinosa polisaha-

rida ekstrakcijom sa toplom vodom, odnosno kiselim rastvorom, korišćena je metodologija

odgovora površine, Box–Behnken dizajnu sa tri nivo i 3, odnosno 4 faktora. Eksperimen-

talni podaci su fitovani polinomnom jednačinom drugog reda i analizirani odgovarajucim

statističkim metodama, pri čemu su određeni odgovarajuci optimalni uslovi obadve eks-

trakcione metode. Pod optimalnim uslovima, eksperimentalni prinos se dobro slaže sa

vrednostima koje su izračunate modelom. Prinos polisaharida dobijenih kiselom ekstrak-

cijom iznosi 42,13 mg·g-1 i veći je od prinosa ostvarenog ekstrakcijom pomoću tople vode

(20,07 mg·g-1). Obe kvadratne jednačine se mogu koristiti za predviđanje rezultata eks-

trakcija.

Ključne reči: proso, polisaharidi, ekstrakcija toplom vodom, kisela ekstrakcija,

metodologija odgovora površine.

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ISSN 1451 - 9372(Print) ISSN 2217 - 7434(Online) JANUARY ... No1_p1-150_Jan-M… · Djordjevic, Modelling of polyester fabric dyeing in the presence of ultrasonic waves ..... 131