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Civil Engineering

Journal (ISSN: 2476-3055)

Editor in Chief:

Prof. M. R. Kavianpour

K.N.Toosi University of Technology (Iran)

Executive Manager:

Dr. O. Aminoroayaie Yamini

K.N.Toosi University of Technology (Iran)

Dr. S. Hooman Mousavi

K.N.Toosi University of Technology (Iran)

Editorial Board Members:

Prof. Dintie S. Mahamah

St. Martin's University (USA)

Dr. Kartik Venkataraman

Tarleton State University (USA)

Dr. Tanya Igneva

University of ACEG (Bulgaria)

Dr. Daniele Bocchiol

Polytechnic University of Milan (Italy)

Dr. Michele Iervolino

Second University of Naples (Italy)

Dr. Rouzbeh Nazari

Rowan University (USA)

Prof. Marta Bottero

Polytechnic University of Turin (Italy)

Chris A. O’Riordan-Adjah (PhD Candidate)

University of Central Florida (USA)

Dr. Yasser Khodair

Bradley University (USA)

Dr. Weidong Wu

University of Tennessee - Chattanooga (USA)

Dr. Kaveh Saleh

University of Sherbrooke (Canada)

To view all editorial board members Click Hear.

Mailing Address: Dr. Kavianpour office, 3rd Floor of Civil Engineering Faculty, K.N.Toosi University of Technology, No. 1346, ValiAsr St, Mirdamad Intersection, Tehran, Iran

Phone: +98-21-88779475-ext. 258 Fax: . +98-21-88779674 E-mail: [email protected] Website: www.CivileJournal.org

Dr. Jiliang Li

Purdue University North Central (USA)

Dr. Yaqi Wanyan

Texas Southern University (USA)

Prof. M.M. Rashidi

Tongji University (China)

Dr. Sanjay Tewari

Louisiana Tech University (USA)

Prof. Nikolaos Eliou

University of Thessaly (Greece)

Dr. Mohammad Reza Najafi

University of Victoria (Canada)

Dr. Saeed Khorram

Eastern Mediterranean University (Cyprus)

Dr. Xinqun Zhu

University of Western Sydney (Australia)

Dr. Jalil Kianfar

St. Louis University (USA)

Dr. Luca Comegna

Second University of Naples (Italy)

Dr. Davide Dalmazzo

Polytechnic University of Turin (Italy)

Dr. Viviana Letelier González

University of the Frontera (Chile)

Dr. Paola Antonaci

Polytechnic University of Turin (Italy)

Dr. Davorin Penava

University of Osijek (Croatia)

Dr. Ali Behnood

Purdue University (USA)

Dr. Uğur Albayrak

Eskisehir Osmangazi University (Turkey)

5th Issue

www.CivileJournal.org

2020

Civil Engineering

Journal (ISSN: 2476-3055)

Contents Page 848-859

On the Characteristics of Ground Motion and the Improvement of the Input Mode of Complex Layered Sites

Hongke Pan, Xinxin Jiang

Page 860-876

Investigating the Flow Hydrodynamics in a Compound Channel with Layered Vegetated Floodplains

Muhammad Ahmad, Usman Ghani, Naveed Anjum, Ghufran Ahmed Pasha, Muhammad Kaleem Ullah, Afzal Ahmed

Page 877-888

The Effects of Nano Bentonite and Fatty Arbocel on Improving the Behavior of Warm Mixture Asphalt against Mois-

ture Damage and Rutting

Sepehr Saedi, Seref Oruc

Page 889-906

Post-Fire Behavior of Post-Tensioned Segmental Concrete Beams under Monotonic Static Loading

Nazar Oukaili, Amer F. Izzet, Haider M. Hekmet

Page 907-918

Optimalization of the Ferronickel Production Process through Improving Desulfurization Effectiveness

Izet Ibrahimi, Nurten Deva, Sabri Mehmeti

Page 919-927

Development of Filters with Minimal Hydraulic Resistance for Underground Water Intakes

A. A. Akulshin, N. V. Bredikhina, An. A. Akulshin, I. Y. Aksenteva, N. P. Ermakova

Page 928-944

Ranking and Determining the Factors Affecting the Road Freight Accidents Model

Masoud Bagheri Ramiani, Gholamreza Shirazian

Page 945-953

Free Vibration of Tall Buildings using Energy Method and Hamilton’s Principle

Peyman Rahgozar

Page 954-960

Smart City and Modelling of Its Unorganized Flows Using Cell Machines

Truong Thanh Trung

Page 961-973

The Effects of Different Shaped Baffle Blocks on the Energy Dissipation

Nassrin Jassim Hussien Al-Mansori, Thair Jabbar Mizhir Alfatlawi, Khalid S. Hashim, Laith S. Al-Zubaidi

Vol. 6, No. 5, May, 2020

www.CivileJournal.org

5th Issue 2020

Civil Engineering

Journal (ISSN: 2476-3055)

Contents Page 974-996

Numerical Investigation of Stress Block for High Strength Concrete Columns

Nizar Assi, Husain Al-Gahtani, Mohammed A. Al-Osta

Page 997-1006

Rational Organizational and Technological Solutions for Plastering

Boris Vasilyevich Zhadanovsky, Vladimir Evgenievich Bazanov

Page 1007-1016

Microstructural and Compressive Strength Analysis for Cement Mortar with Industrial Waste Materials

Zahraa Fakhri Jawad, Rusul Jaber Ghayyib, Awham Jumah Salman

Page 1017-1030

Improving the Aging Resistance of Asphalt by Addition of Polyethylene and Sulphur

Maria Iqbal, Arshad Hussain, Afaq Khattak, Kamran Ahmad

Page 1031-1038

Mechanical Properties of Cement Mortar after Dry–Wet Cycles and High Temperature

Xiong Liang-Xiao, Song Xiao-Gang

www.CivileJournal.org

5th Issue 2020

Vol. 6, No. 5, May, 2020

Focus and Scope

Civil Engineering Journal (C.E.J) is a multidisciplinary, an open-access, internationally double-blind peer-reviewed journal concerned with all aspects of civil engineering, which include but are not necessarily restricted to:

Special Issues

Special Issues deal with more focused topics with high current interest falling within the scope of the journal in which they are published. Special Issue proposals are welcome at any time during the year.

For most of the civil engineering conferences it is possible to submit papers presented at the conference for subsequent publication in special issues of the C.E.J.

Civil Engineering Journal (C.E.J) is published monthly.

Civil Engineering Journal (C.E.J) has fast peer review process (3-4 weeks).

Civil Engineering Journal (C.E.J) Indexing & Abstracting

This is an open access journal under the CC-

BY license (https:// crea-tivecommons.org/licenses/ by/4.0/).

Civ

il E

ngi

nee

rin

g Jo

urn

al

I

• Building Materials and Structures • Coastal and Harbor Engineering

• Constructions Technology • Constructions Economy and Management

• Earthquake Engineering • Environmental Engineering

• Renovation of Buildings • Geotechnical Engineering

• Highway Engineering • Hydraulic and Hydraulic Structures

• Road and Bridge Engineering • Structural Engineering

• Surveying and Geo-Spatial Engineering • Transportation Engineering

• Tunnel Engineering • Urban Engineering and Economy

• Water Resources Engineering • Urban Drainage

Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

848

On the Characteristics of Ground Motion and the Improvement

of the Input Mode of Complex Layered Sites

Hongke Pan a, Xinxin Jiang

a, b*

a School of Building Engineering, Xinyu University, Xinyu City 338004, China.

b Earthquake Engineering Research Center, China Institute of Water Resources and Hydropower Research, Beijing 100038, China.

Received 28 January 2020; Accepted 05 April 2020

Abstract

It is a hot research topic to perform the dynamic interaction analysis between the engineering structure and the soil by

using the time-domain method. This paper studies the seismic behaviour of the layered sites and the seismic response of

the structures using the viscous-spring artificial boundary theory. The artificial boundary model of viscous-spring is

initially based on homogeneous foundation. For the layered site (Foundation), the traditional homogeneous model or

equivalent load input mode is not suitable, which may bring great error. By introducing the changes of coefficients and

phases of reflection and transmission of seismic waves at the interface between layers, an improved method of equivalent

load input mode of traditional viscous-spring artificial boundary model is proposed. This new wave model can simulate

the propagation law of seismic wave in layered site more accurately, which is available for the seismic performance of

engineering structure under the condition of large and complex layered site. At last, the simplified homogeneous model,

the equivalent load input method and the improved layered model input method are used to study the seismic response of

the engineering example. It is shown that the results calculated by the three methods are different, which shows that the

homogeneous foundation model and the conventional equivalent load input method of seismic wave cannot simulate the

seismic force accurately, whereas the improved wave input model can better reflect the characteristic of traveling wave

in layered sites.

Keywords: Earthquake Resistance of Engineering Structures; Layered Foundation; Time Domain Analysis Method; Seismic Wave

Propagation; Improvement of Input Mode.

1. Introduction

Complex layered sites are often encountered in the construction process of various large-scale civil, water

conservancy and transportation projects. Although the effects of layered sites on the structural dynamic response have

been acknowledged, there is no comprehensive understanding and design experience of it due to the complexities.

Consequently, how to evaluate the dynamic response characteristics, seismic stability and seismic measures of the

interaction between superstructure and soil has become a difficult problem for the builders, and sometimes even

directly affects the construction of the project. This paper is devoted to the analysis of structural dynamic response in

layered sites.

It has been acknowledged that the dynamic soil-structure interaction in the analysis of seismic response should be

considered. Various numerical methods are developed to simulate the seismic response, and they can be inducted as

global artificial boundary and local artificial boundary theory [1]. The global artificial boundary theory, typically

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091512

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

849

represented by boundary element method [2] and scaled boundary finite element method [3], can satisfy all field

formulas and boundary conditions in infinite domain. The local artificial boundary theory is developed on the concept

of unilateral wave, which is typical represented by viscous boundary method [4], viscous-spring artificial boundary

method [5] and transmission boundary method [6]. These initial patterns of the dynamic soil-structure interaction are

established with homogenous soil. Afterwards, in order to deal with more complex engineering conditions, the

researches carry out experiments on complex sites and propose various methods to broaden its application. For

example, Moghadam and Baziar investigated the effect of a circular subway tunnel on the ground motion

amplification pattern by Shaking table testing and numerical simulation [7]. Sun et al. used the analytical study and

numerical analysis to characterize the underlying soft soil layer-tunnel interaction problem [8]. Karabalis and

Mohammadi used a 3-D frequency domain BEM equations in conjunction with infinite space fundamental solutions to

simulate the layered soil medium [9]. Birk and Behnke derived a modified SBFEM for the analysis of 3D-layered

continua based on the use of a scaling line instead of a scaling centre, and the dynamic stiffness coefficients were

calculated to demonstrate the accuracy of the method [10]. Li et al. developed a time-domain method to calculate the

free field motion of a layered half-space subjected to oblique incident body waves [11]. Liu and Wang developed a 1D

finite element method in time domain to calculate the in-plane wave motion of free field in elastic layered space by

oblique seismic incidence [12].

At present, the analysis of structure-foundation interaction using time-domain method is a hot topic, in which the

viscous-spring artificial boundary performs well in the application of computational precision, stability and large-scale

finite element software, and its ground motion input mode is easy to simulate the propagation process of seismic wave

and the non-uniform change of field motion caused by the oblique incidence of seismic wave, so it has been widely

used [12]. The viscous-spring artificial boundary theory is originally based on the homogenous soil. When it comes to

complex layered sites, the major challenge is to simulate the seismic wave in the truncation boundary. It is well known

that the seismic wave will be reflected and transmitted at the soil interface due to the different mechanical parameters

of the soil. However, most of the existing research and engineering applications do not examine that characteristic of

layered sites. One method is to take the soil parameters adjacent to the foundation as homogeneous soil model instead

of the layered sites. Another method is based on layered foundation model but adopt conventional equivalent load

mode of ground motion input mode, which only converts incident wave and reflected wave at the free top surface into

equivalent load form and then inputs them into the system to solve the dynamic response of the whole system. For

homogeneous sites, seismic waves are reflected at the free surface, and the equivalent load reflects the superposition

effect of incident waves and reflected waves. For layered foundations, seismic waves will reflect and transmit at the

interface of the interlayer materials, and the amplitude and phase of the reflected and transmitted waves will change

[13-15]. If the conventional equivalent load input mode is still used in layered sites, it is difficult to consider the

variation of the amplitude and phase of the fluctuation, which makes the fluctuation input of the artificial boundary

node different from the actual fluctuation amplitude and phase of the internal node of the layered foundation model,

resulting in the inconsistency of the vibration of the boundary node and the internal node of the model, thus causing

relatively large calculation errors [16-18].

In order to simulate the propagation of seismic waves in a layered foundation, the input model of seismic wave

under equivalent load is improved based on the viscous-spring artificial boundary model, and the reflection and

transmission coefficient and phase change values of seismic waves at the material interface are introduced, so as to

improve the equivalent load input mode for wave input method under layered foundation. The method inherits the

advantages of viscous-spring artificial boundary.

This paper commences with the wave propagation process in a two layered site. The phenomenon of seismic

reflection and transmission at the interface of the soil is illustrated, followed by derivation of formulas for reflection

and transmission of seismic waves at the interface between layers. And an improved input mode of seismic waves in

layered sites is proposed based on these formulas. Engineering example demonstrates the accuracy of the new seismic

input mode and its implementation.

2. Wave Characteristics of Seismic Waves in Layered Sites

2.1. Multiple Reflection and Transmission of Seismic Waves at Horizontal Interfaces

The change of amplitude and phase of seismic wave propagating in layered site (foundation), showing different

propagating characteristics compared with that in homogeneous foundation, is mainly caused by the different material

parameters of each layer. The change of mechanical parameters at the interlayer interface makes part of the wave

propagation energy reflect back to the lower soil layer, while the other part of the wave energy continues to propagate

upward. Figure 1 shows the propagation process of the seismic wave vertically incident to the two-layer half-space

free field. The interfaces are numbered 0, 1, and 2 from bottom to top, the interface 1 is interlayer material interface,

and the interface 2 is free surface, the shear wave velocities of the first and second layers of soil are Cs1 and Cs2

respectively, the initial input time of seismic wave is 0. The wave pattern of typical time is selected, and the arrow is

the wave front of seismic wave.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

850

1) When Δ𝑡 =ℎ1

𝐶𝑠1: the seismic wave is reflected and transmitted at interface 1, and the amplitudes of incident wave,

reflected wave and transmitted wave are recorded as 𝐴𝑖1, 𝐴𝑟1 and 𝐴𝑡1 respectively;

2) When Δ𝑡 =ℎ1

𝐶𝑠1+

ℎ2

𝐶𝑠2: the seismic wave is reflected at interface 2, and 𝐴𝑖1 , 𝐴𝑟2 represent the amplitude of the

incident wave and the reflected wave;

3) When Δ𝑡 =ℎ1

𝐶𝑠1+

2ℎ2

𝐶𝑠2, two seismic waves are reflected and transmitted simultaneously at the interface 1. Firstly, the

downward seismic waves are reflected by the top surface, 𝐴𝑖3, 𝐴𝑟3 and 𝐴𝑡3 are the amplitudes of incident wave,

reflected wave and transmitted wave respectively, and 𝐴𝑖3 = 𝐴𝑟2 . Secondly, the upward seismic waves are

reflected by the ground, similarly, 𝐴′𝑖3 , 𝐴′𝑟3and 𝐴′𝑡3 are the amplitudes of incident wave, reflected wave and

transmitted wave respectively;

4) Δ𝑡 =ℎ1

𝐶𝑠1+

3ℎ2

𝐶𝑠2: two waves propagating upward on the top surface with the amplitude 𝐴𝑟3 and 𝐴′𝑡3 are reflected

simultaneously. The phases change after reflection;

5) As shown in the previous four steps, the seismic waves will be reflected and transmitted continuously at interface 1

and 2 during the propagation process. The seismic waves in the foundation are superimposed on each other, and

some wave forms will persist in the soil layer.

2.2. Equations for Reflection and Transmission of Seismic Waves at the Interface between Layers

As shown in Figure 2, the seismic wave is obliquely incident at the horizontal interface, and reflection and

transmission occur. Snell's theorem states that the various waves in the wave system on the interface have the same

apparent propagation velocity along the interface, and the corresponding mathematical expression is expressed as:

1 1 2

sin sin sinish rsh tsh

s s sC C C

(1)

With

ish rsh (2)

The incident wave, reflected wave and transmitted wave in Figure 2 are respectively represented by subscripts i, r

and t; A represents the amplitude of the wave; θ is the angle between the direction of wave propagation and the

interface normal.

Figure 1. The sketch of seismic wave reflection and transmission on the interface

A ' i3

Seismic wave continues

to propagate

A i4

The first layer:

shear wave velocity C s1

A ' i4

1.Time t=h 1 /C s1 ,the wave front

is at interface 1

3.Time t=h 1 /C s1 +2h 2 /C s2 , the wave front

is at interface 1

A t1

A i2

A ' r3

A r4

4.Time t=h 1 /C s1 +3h 2 /C s2 , the wave

front is at interface 2

The second layer:

shear wave velocity C s2

A r1

......

A i

A r2

A i3

h 2

Interface 1 A ' t3

A ' r4

A t3

h 1

Interface 0

A r3

Interface 2

0.Time t=0.The wave shoot in from the bottom

2.Time t=h 1 /C s1 +h 2 /C s2 , the wave front

is at interface 2

A i1

Civil Engineering Journal Vol. 6, No. 5, May, 2020

851

Figure 2. The sketch map of reflection and transmission of the SH wave on the interface

The relationship between the amplitude of incident SH wave, reflected SH wave and transmitted SH wave is

derived below. Assuming that the material is in close contact at the interface, the equations for the boundary

conditions at the interface 1 are obtained by the compatibility and equilibrium at the interface.

(1) (2)U U (3)

(1) (2)

xy xy (4)

Where U is the horizontal displacement; 𝜏 is the shear stress; the superscript 1 and 2 denote the lower soil layer 1 and

the upper soil layer 2, respectively.

𝑈𝑖𝑠ℎ , 𝑈𝑟𝑠ℎ and 𝑈𝑡𝑠ℎ are assumed respectively (only Y-direction deformation occurs).

1

cosexp[ ( )]ish

ish ish

s

U A i t yC

(5)

1

cosexp[ ( )]rsh

rsh rsh

s

U A i t yC

(6)

2

cosexp[ ( )]tsh

tsh tsh

s

U A i t yC

(7)

The displacements of soil layer 1 and 2 in Figure 2 are respectively:

(1) (1) (1)

ish rshU U U (8)

(2)

tshU U (9)

In the above formulas, 𝜏𝑦𝑥(1)

, 𝜏𝑦𝑥(2)

can be derived from the linear elastic stress-strain relation 𝜏 = 𝐺𝜕𝑢

𝜕𝑦. Substituting

Equations 8 and 9 and the derived 𝜏𝑦𝑥(1)

, 𝜏𝑦𝑥(2)

into the boundary conditions of Equations 3 and 4, yields:

1 rsh tsh

ish ish

A A

A A (10)

1 1 1 1 2 2cos cos cos cosrsh tshs isH s rsh s tsh tsH

ish ish

A AC C C

A A

(11)

Substituting Equation 10 into 11, and noting that 𝜃𝑖𝑠ℎ = 𝜃𝑟𝑠ℎ, the seismic reflection coefficient and transmission

coefficient formulas are expressed as:

1 1 2 2

1 1 2 2

cos cos

cos cos

rsh s ish s tsh

ish s ish s tsh

A C C

A C C

(12)

1 1

1 1 2 2

2 cos

cos cos

tsh s ish

ish s ish s tsh

A C

A C C

(13)

X

physical and mechanicalparametersρ2,Cs2,μ2

θrsh

Y

physical and mechanicalparametersρ1,Cs1,μ1

incident SH wave

upper layer 2

θish

θtsh

Aish Arsh

Atshtransmitted SH wave

lower layer 1

reflected SH wave

Civil Engineering Journal Vol. 6, No. 5, May, 2020

852

Typically, when the incident wave is perpendicular to the interface, 𝜃𝑖𝑠ℎ = 90𝑜 . Substituting 𝛼 =𝜌2𝐶𝑠2

𝜌1𝐶𝑠1 into

Equations 12 and 13, they are simplified as:

1

1rsh

ish

A

A

(14)

2

1tsh

ish

A

A

(15)

If the interface is a free surface, then: 𝛼 = 0,𝐴𝑟𝑠ℎ

𝐴𝑖𝑠ℎ= 1.

P wave and SV wave can be derived similarly when they are incident, but it should be noted that there will be

waveform conversion at the interface of P wave and SV wave oblique incidence, that is, SV wave and P wave exist in

both reflected wave and transmitted wave.

3. Input Mode of Seismic Waves in Layered Sites

3.1. Improved Wave Input Mode of Equivalent Load Form

The conventional expression of the equivalent load of the traditional viscous-spring artificial boundary is express

as [18]:

𝐹𝐵(𝑡) = 𝜏0(𝑥𝐵 , 𝑦𝐵 , 𝑡) + 𝐶𝐵𝜔0(𝑥𝐵 , 𝑦𝐵 , 𝑡) + 𝐾𝐵𝜔0(𝑥𝐵 , 𝑦𝐵 , 𝑡) (16)

Where, 𝑥𝐵 , 𝑦𝐵 and t are the coordinates and time of point B of the artificial boundary successively;

𝜔0(𝑥𝐵 , 𝑦𝐵 , 𝑡) stands for the displacement and velocity of the incident wave field at the node; 𝐶𝐵 and 𝐾𝐵 are the damper

coefficient and the spring elasticity coefficient, respectively.

The seismic wave of layered foundation is reflected and transmitted at the interface between layers, and the

displacement amplitude and phase change. Displacement 𝜔𝑟(𝑥𝐵 , 𝑦𝐵 , 𝑡) of the reflected wave and displacement

𝜔𝑡(𝑥𝐵 , 𝑦𝐵 , 𝑡) of the transmitted wave are expressed as:

0( , , ) ( , , )rr B B B B

i

Ax y t x y t

A (17)

0( , , ) ( , , )tt B B B B

i

Ax y t x y t

A

(18)

After derivation of Equations 17 and 18, the corresponding expression of node velocity can be obtained, and then

substituting it into Equation 16, the expression 𝐹′𝐵(𝑡) of the improved equivalent load of reflection and transmission

of seismic wave at the interface is obtained:

𝐹′𝐵(𝑡) = 𝜏0(𝑥𝐵 , 𝑦𝐵 , 𝑡) + 𝐶𝐵 [𝐴𝑡

𝐴𝑖

𝜔0(𝑥𝐵 , 𝑦𝐵 , 𝑡) −𝐴𝑟

𝐴𝑖

𝜔0(𝑥𝐵 , 𝑦𝐵 , 𝑡)] + 𝐾𝐵 [𝐴𝑡

𝐴𝑖

𝜔0(𝑥𝐵 , 𝑦𝐵 , 𝑡) +𝐴𝑟

𝐴𝑖

𝜔0(𝑥𝐵 , 𝑦𝐵 , 𝑡)] (19)

If the physical and mechanical parameters of the upper and lower layers of the interface are the same, the layered

foundation will be degraded into a homogeneous site. 𝐴𝑡

𝐴𝑖= 1,

𝐴𝑟

𝐴𝑖= 0. It means that seismic waves at interlayer

interface do not reflect, and the improved expression (19) of equivalent load can be changed to (16), which indirectly

verifies the correctness of the derivation process.

3.2. Realization of Wave Input of Layered Site in ANSYS

In Software ANSYS, the wave input of viscous-spring artificial boundary of layered foundation is redeveloped,

and then the macro file is made to generate ANSYS tool button, which can automatically load the input file of seismic

wave and the viscous-spring artificial boundary for seismic response analysis. It includes spring damper application

module, equivalent load generation and loading module, solver solution and post-processing module.

1) Automatic application of spring damper: The boundary of the foundation is a box-shaped artificial boundary, which

can automatically search for the boundary nodes and store them. The spring damper of the boundary nodes can be

applied automatically by “Ndnext” command.

2) Generation and application of equivalent load: In the process of simulating multiple reflection and transmission of

seismic waves, two parameters of two-dimensional array can be used to track the position of seismic wave front

and the number of times of transmission and reflection respectively, so as to express the delay phase of seismic

Civil Engineering Journal Vol. 6, No. 5, May, 2020

853

waves and the amplitude after multiple reflection and transmission. For each solution time, the equivalent load on

the node after each wave superposition is calculated and applied to the boundary node for solution, which ensures

the accurate simulation of wave input at each time.

3) Solver solution and post-processing: The Newmark integral is used to solve the dynamic response in ANSYS

transient analysis, and linear elastic solution is used in this paper. Post-processing mainly uses Post1 and Post26 to

extract stress, displacement, acceleration and draw contours.

4. Engineering Example

A roller compacted concrete gravity dam located on two layers of foundation is chosen to illustrate actual

engineering situation that can be modelled accurately by applying the improved seismic input mode. The analysis of

the free field of layered site showed that the seismic behaviour of the layered foundation was significantly different

from the homogenous foundation [12]. Next, the engineering adaptability of the gravity dam model is verified by the

seismic response analysis of the model applied to the two-story site.

4.1. Basic Information

The dam is a roller compacted concrete gravity dam located on two layers of foundation. The height of the dam

body is 73 m, the top thickness of the dam is 6 m, and the slope of the downstream dam is 1:0.7. There are two types

of materials for dam concrete: the dynamic elastic modulus of concrete outside the dam is 36.4 GPa, the density is

2400 kg/m3, and the Poisson ratio is 0.167; while the dynamic elastic modulus of concrete inside the dam is 28.6 GPa,

the density is 2400 kg/m3, and the Poisson ratio is 0.167. There are two types of materials for foundation from bottom

to top: the dynamic elastic modulus at the bottom is 52 GPa, density is 2700 kg/m3, Poisson ratio is 0.25; while the

dynamic elastic modulus at the top is 26 GPa, density is 2700 kg/m3, Poisson ratio is 0.25. The designed peak ground

motion acceleration at the dam site is 0.374 g.

4.2. Finite Element Model of Gravity Dam-Layered Site

In the finite element model of gravity dam-layered site shown in Figure 3, the dam body is divided into grids by

Plane42 unit, which is set to the plane strain problem. A total of 3164 units and 7778 nodes are divided. The viscous-

spring artificial boundary is simulated by spring damper, which is applied automatically by node number. The

scattering source is located at the centroid of the gravity dam, and the spring coefficient R of each boundary spring

damper is taken as the average value of the distance from the scattering source to each boundary surface and to the

corner of the boundary. The spring damper coefficient at the material interface is taken as the average value of the

upper and lower layers.

Figure 3. The finite element model of gravity dam and foundation

4.3. Seismic Wave Processing

The designed peak ground motion acceleration of the site where the dam is located is 0.374 g. The target response

spectrum is determined by referring to the Specifications for Seismic Design of Hydraulic Structures DL5073-2000

[19], and then the artificial seismic wave is synthesized. Wherein, the representative value 𝛽𝑚𝑎𝑥 of the maximum

design response spectrum is 2.0 for gravity dams and 20% for the lower limit of design response spectrum, i.e. 0.4 for

maximum representative value 𝛽𝑚𝑖𝑛 , the site should be determined as category 1 according to the foundation

Civil Engineering Journal Vol. 6, No. 5, May, 2020

854

parameters, and the characteristic period 𝑇𝑔 is 0.20 s, thus the design response spectrum is determined, and the

artificial seismic wave is synthesized according to the response spectrum, and the duration of the artificial seismic

wave is 28 s, as shown in Figure 4.

A. Comparison between calculation response spectrum and design response spectrum

B. Acceleration time history of artificial seismic waves

C. Velocity time history of artificial seismic waves

0 1 2 3

0.5

1.0

1.5

2.0

Response spectrum value

T/s

The response spectrum calculated by artificial wave

Design response spectrum

0 4 8 12 16 20 24 28

-1.0

-0.5

0.0

0.5

1.0

Artificial seismic wave

Acc

eler

atio

n(m

/s2)

Time(s)

0 4 8 12 16 20 24 28

-0.3

-0.2

-0.1

0.0

0.1

0.2

Vel

oci

ty(m

/s)

Time(s)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

855

D. Displacement time history of artificial seismic waves

Figure 4. The Artificial Seismic Wave

4.4. Calculation Conditions

The seismic response calculation method considering layered foundation structure at traditional viscous-spring

artificial boundary is compared with the viscous-spring artificial boundary model and wave input method considering

seismic wave reflection and transmission at material interface in this paper, and three calculation conditions are

designed respectively, as shown in Table 1. Only the horizontal shear wave of seismic wave is input to calculate the

response of empty reservoir without water under the horizontal earthquake.

Table 1. The model and mechanical parameters of three methods

Scheme 1 (Case 1) Scheme 2 Scheme 3

Model description Simplified homogeneous model Layered site model (Equivalent load inp

ut mode of seismic wave)

Layered site model (Improved input mode

of seismic wave)

Stratum parameter Foundation density is 2700 kg/m3, elastic modulus is 26 GPa, Poisson

ratio is 0.167.

The density of upper foundation is 2700

kg/m3, the elastic modulus is 26 GPa, and Poisson ratio is 0.167; the density of th

e lower foundation is 2700kg/m3, the elastic modulus is 52 GPa, and the Poisson

ratio is 0.167.

The density of upper foundation is 2700 k

g/m3, the elastic modulus is 26 GPa, and Poisson ratio is 0.167; the density of the lo

wer foundation is 2700 kg/m3, the elastic modulus is 52 GPa, and the Poisson ratio i

s 0.167.

Model characteristics The homogeneous foundation has no reflection and transmission of s

eismic waves at the interface.

The wave input does not take into account the reflection and transmission of seis

mic waves at the interface.

The wave input takes into account the reflection and transmission of seismic waves

at the interface.

4.5. Arrangement of Calculation Results

4.5.1. Calculation results of displacement and acceleration

The results of displacement and acceleration calculation are shown in Figure 5.

0 4 8 12 16 20 24 28

-0.3

-0.2

-0.1

0.0

0.1

0.2

0.3

Dis

pla

ce

me

nt(

m)

Time(s)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

856

A. Contrast diagram of dam top displacement under three working conditions

B. Contrast diagram of acceleration of dam top under three working conditions

Figure 5. The top results of displacement and acceleration

4.5.2. Calculation Results of Stress Diagram

The stress diagrams are shown in Figures 6 to 8.

0 4 8 12 16 20 24 28

-2.0

-1.5

-1.0

-0.5

0.0

0.5

1.0

1.5

2.0

Dis

pla

ce

me

nt(

m)

Time(s)

Case 1

Case 2

Case 3

0 4 8 12 16 20 24 28

-20

0

20

Acce

lera

tio

n(m

/s2)

Time(s)

case1

case2

case3

A. Isogram of local vertical stress in dam heel B. Isogram of local first principal stress at dam heel

Figure 6. The results of the scheme 1

ANSYS 13.0

B

A

B

A

A

B

A

B

A

B

C

D

B

C

B

E H

IFG

D C

B

A

C

BA

DE

C

B

A

A =104211

B =312634

C =521056

D =729479

E =937901

F =.115E+07

G =.135E+07

H =.156E+07

I =.177E+07

ANSYS 13.0

F

E

D

C

CD

G

H

F

C

C

G

H

E

F

GH

D

C

C

E F D

C

A

B

C

A

E

B

D

A

A =336118

B =983299

C =.163E+07

D =.228E+07

E =.292E+07

F =.357E+07

G =.422E+07

H =.487E+07

I =.551E+07

Civil Engineering Journal Vol. 6, No. 5, May, 2020

857

4.6. Results and Discussion

Figure 5 plots the displacement and acceleration time history of the extracted dam crest joints. The displacement

amplitude in scheme one is 1.37 m, and that of scheme two is approximately the same as that of scheme three, which

is 1.74m, indicating that the calculation result of layered foundation is smaller when it is simplified into homogeneous

foundation. The results of acceleration calculation also illustrate this problem.

The exact stress distribution of the local stress in the heel of the dam under the three schemes is also plotted in

Figures 6 to 8. The local vertical stress distribution is basically the same as the isogram of the first principal stress, but

the values are obviously different, which are successively scheme 3, scheme 2 and scheme 1 from the largest to the

smallest. The stresses in the heel of the dam of scheme 2 differ greatly from that of scheme 3, in which the relative

error of vertical stress is 37.2%, and the relative error of principal stress is 10.5%. Combined with Figures 6 to 8, it

can be seen that although the displacement and acceleration of the dam crest calculated in Scheme 2 are similar to

those in Scheme 3, the significant difference in the results of heel stress calculation illustrates the necessity of

considering the wave input method of reflection and transmission of seismic wave in Scheme 3, because the wave

input of boundary node is inconsistent with the vibration of internal node in Scheme 2, which restricts the propagation

of real seismic wave. At the same time, the calculation results show that, for such gravity concrete dam, the stress at

the heel of gravity dam is the key part of earthquake resistance, which needs to be paid attention to in the seismic

design and construction of gravity dam.

In this example, the presented results show that: for the layered site, when the physical and mechanical parameters

of the soil layers are quite different, the model of simplified homogeneous foundation may bring great errors;

similarly, when the layered foundation model is adopted, the conventional equivalent load input mode of seismic wave

also cause errors compared with the improved input mode of seismic wave. The traditional equivalent load input mode

A. Isogram of local vertical stress in dam heel B. Isogram of local first principal stress at dam heel

Figure 7. The results of the scheme 2

A. Isogram of local vertical stress in dam heel B. Isogram of local first principal stress at dam heel

Figure 8. The results of the scheme 3

ANSYS 12.1

B

A

A

BA

A

B

A

B

C

B

D

C

B

B

A

E

G

H

F

D

C

B

C

A

B

CDE

AB

A =142148

B =426443

C =710739

D =995035

E =.128E+07

F =.156E+07

G =.185E+07

H =.213E+07

I =.242E+07

ANSYS 12.1

F

ED

C

E D

CD

G

F

G

C

C

H

E

F

G

H

D

C

C

E F

D C

A

B

C

A

DE

B

A

A =490862

B =.143E+07

C =.237E+07

D =.332E+07

E =.426E+07

F =.520E+07

G =.614E+07

H =.708E+07

I =.803E+07

ANSYS 13.0

A

B

A

A

A

B

A

B

A

B

C

D

B

B

B

AC

EF

HI

G

D

CB

ABC

DE

A

F

B

C

A =195514

B =586543

C =977572

D =.137E+07

E =.176E+07

F =.215E+07

G =.254E+07

H =.293E+07

I =.332E+07

ANSYS 13.0

F

E

D

G

C

E

D

F

G

C

G

F

H

E

D

C

C

FG

E

D

C

BAB

C

D

A

A =541073

B =.158E+07

C =.262E+07

D =.366E+07

E =.471E+07

F =.575E+07

G =.679E+07

H =.783E+07

I =.887E+07

Civil Engineering Journal Vol. 6, No. 5, May, 2020

858

is mainly applicable to the case of homogeneous foundation at first, which only considers the superposition effect of

incident wave and reflected wave of seismic wave at top free surface. However, in the actual layered foundation,

seismic wave will be reflected and transmitted at the interface between layers (and may occur continuous reflection

and transmission phenomenon). Therefore, this paper proposes to consider the reflection and transmission coefficients

and phase changes of seismic wave at the interlayer interface, so as to reflect the actual propagation law and mode of

seismic wave.

5. Conclusion

This paper study the seismic behavior of layered sites and has developed an improved seismic input mode of the

viscous-spring artificial boundary theory based on the characteristic of traveling wave. For the seismic response

analysis of layered site, its dynamic characteristics have an important influence on the structural dynamic response.

When viscous-spring artificial boundary is used to simulate the dynamic interaction between the structure and the

foundation, the homogeneous foundation model is difficult to reflect the propagation characteristics of seismic waves

between the soil layers. At present, the most widely used method is the layered site model - the equivalent load input

mode of seismic waves. However, engineering experience and theoretical analysis show that the traditional wave input

mode used in layered site condition may lead to the inconsistency between the wave input of the boundary node and

the vibration of the internal node, thus causing the calculation error. Considering the reflection and transmission of

seismic wave on the layered interface, the improved method can better reflect the actual situation of the site and

improve the accuracy of the calculation results.

Compared with the conventional equivalent load input mode, the improved wave input mode continuously tracks

the propagation process of seismic wave, keeps the consistency between the boundary input and the internal node

vibration, thus expanding the application scope of the original viscous-spring artificial boundary model. In addition, in

the author's further study; it has been preliminarily found that whether it is necessary to adopt the improved seismic

wave input mode method for layered foundation is also closely related to the physical quantities such as the shear

modulus of the stratum and the density of soil. Only when these physical quantities between the soil layers reach a

certain difference, the improved seismic wave input mode should be considered. The specific difference of these

physical quantities needs further study.

6. Funding

The research described in this paper was financially supported by science and technology research project of

Jiangxi Provincial Education Department (Grant No. GJJ171061).

7. Conflicts of Interest

The authors declare no conflict of interest.

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[4] Miura, Fusanori, and Hiroshi Okinaka. “Dynamic Analysis Method for 3-D Soil-Structure Interaction Systems with the Viscous

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[5] Wang, Duguo, Peixin Shi, and Chenggang Zhao. “Two-Dimensional in-Plane Seismic Response of Long-Span Bridges Under

Oblique P-Wave Incidence.” Bulletin of Earthquake Engineering 17, no. 9 (June 13, 2019): 5073–5099. doi:10.1007/s10518-

019-00664-7.

[6] Aydinoğlu, M.Nuray. “Consistent Formulation of Direct and Substructure Methods in Nonlinear Soil-Structure Interaction.”

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[7] Rabeti Moghadam, Masoud, and Mohammad Hassan Baziar. “Seismic Ground Motion Amplification Pattern Induced by a

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[8] Sun, Qiangqiang, Daniel Dias, and Luis Ribeiro e Sousa. “Impact of an Underlying Soft Soil Layer on Tunnel Lining in Seismic

Conditions.” Tunnelling and Underground Space Technology 90 (August 2019): 293–308. doi:10.1016/j.tust.2019.05.011.

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Space Subjected to Obliquely Incident Body Waves.” Earthquake Engineering and Engineering Vibration 6, no. 2 (June 2007):

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[13] Pasqua, Fernando Della, Rafael Benites, Chris Massey, and Mauri MacSaveney. “Numerical Evaluation of 2D Versus 3D

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Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

860

Investigating the Flow Hydrodynamics in a Compound Channel

with Layered Vegetated Floodplains

Muhammad Ahmad a*

, Usman Ghani a, Naveed Anjum

b, Ghufran Ahmed Pasha

a,

Muhammad Kaleem Ullah c, Afzal Ahmed

a

a Department of Civil Engineering, University of Engineering and Technology Taxila, Taxila, Pakistan.

b Department of Civil and Environmental Engineering, Graduate School of Science and Engineering, Saitama University, Saitama, Japan.

c Department of Civil Engineering, The University of Lahore, Lahore, Pakistan.

Received 22 December 2019; Accepted 11 April 2020

Abstract

In natural rivers, vegetation grows on floodplains, generating complex velocity field within the compound channel. The

efficient modelling of the flow hydraulics in a compound channel with vegetated floodplains is necessary to understand

and determine the natural processes in rivers and streams. As the three dimensional (3D) flow features are difficult to

capture through experimental investigation; therefore, the present numerical study was carried out to investigate the

complex 3D flow structures with the vertically layered vegetation placed over the floodplains in a symmetric trapezoidal

compound channel. The simulations were conducted using a Computational Fluid Dynamics (CFD) code FLUENT,

whereas a Reynolds Averaged Navier-Stokes (RANS) technique based on Reynolds stress model (RSM) was

implemented for turbulence closure. The numerical model successfully replicated the flow behavior and showed a good

agreement with the experimental data. The present study concluded the presence of quite-S shaped velocity profile in the

layered vegetated floodplains when the short vegetation was submerged during high flows or floods, whereas the

velocity profile was uniform or almost logarithmic during low floods or when both short and tall vegetation remained

emergent. The lateral exchange of mass and momentum was promoted due to the flow separation and instability along

the junction of the floodplains and main channel. The flow velocities were significantly reduced in the floodplains due to

resistance offered by the vegetation, which consequently resulted in an increased percentage i.e. 67-73%, of passing

discharge through the main channel. In general, the spatial distribution of mean flow and turbulence characteristics was

considerably affected near the floodplain and main channel interfaces. Moreover, this study indicated a positive flow

response for the sediment deposition as well as for the nourishment of the aquatic organisms in the riparian environment.

Keywords: Compound Channel; Vegetated Floodplains; Numerical Modelling; Flow Characteristics.

1. Introduction

Vegetation on flood plains is a world-wide engineering problem in most of the natural rivers, which does not only

affect the flow conveyance capacity but also influences the ecological system of rivers [1]. Natural rivers are generally

functioned by the main channel for conveyance of the primary flow and a vegetated floodplain to carry the extra flow

during floods. The vegetation on the floodplain offers hydraulic resistance as it typically leads to the reduced flow

velocity and increases the difference in velocity between the main channel and the floodplain. Many river problems

demand accurate predictions of the flow conveyance in compound channels. It facilitates the engineers in the

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091513

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

861

development of important information related to the flood protection, design of hydraulic structures, and sediment load

estimation to plan for effective prevention measures [2-4]. It is related to other practical engineering problems such as

river training and morphology, dredging and design of flood alleviation works [5-7]. It encourages the hydraulic

engineers to estimate the flood risks and to develop the mitigation schemes [8, 9].

Vegetation on the floodplains has distinct functions to understand the morpho-dynamics and flood control features

of river. It includes flow reduction, streamline regulation and shoreline protection against the erosion. On the other

hand, it enhances sediment deposition and causes overgrowth of the vegetation [1]. Several researchers have explored

different flow patterns and discharge prediction in a vegetated compound channel [10-12]. Flow conveyance in a

compound channel can be accurately determined with considering the momentum-transfer mechanism across the flow

sections [13-15]. Thus, the outcomes of the previous studies identified that the flow structures in a compound channel

are more complex than that in a simple channel.

The flow structure at the interface of main channel and floodplains containing vegetation becomes complex due the

exchange of momentum [16, 17], which affects the overall discharge carrying capacity of the channel [18]. The

vegetated floodplains play a role in sediment trapping due to larger exchange of momentum [19] and helps in channel

restoration [20]. Thus, the introduction of vegetation on the floodplains illustrates the better understanding of flow

hydrodynamics in compound channels.

Previous researchers have investigated the compound channel flow with focusing on constant height of the

vegetation on floodplains under either emergent or submerged flow conditions, which is not as real as that in natural

rivers or streams. In fact, in natural riparian environment, vegetation exists with heterogeneity i.e. short vegetation

such as shrubs and grass, as well as tall vegetation such as trees [21], which experiences both emergent and submerged

conditions. Although there are only a few studies on flows with an array of short and tall vegetation together [21,22] in

rectangular channels; however, the complex interaction of flow in the floodplains having such kind of riparian

vegetation with the main channel flow needs to be explored. Thus, these studies indicate the influence of vegetation

diversity on flow structures in natural riparian environment.

During the higher floods in riparian environment or in flood plains, the short vegetation becomes submerged and

tall vegetation becomes emergent which results in extra complexity of the flow structure due to shear layer formation

around the top of submerged vegetation [1]. On the other hand, when the water level is small during low flows, short

vegetation as well as tall vegetation remains emergent. Thus, it is necessary to consider the effects of the short and tall

vegetation, which contributes an extra part of complexity in the flow patterns, to effectively replicate the riparian

environment. Although understanding of the flow hydrodynamics with incorporating riparian vegetation has

extensively been promoted by the previous studies, discussion is still limited within the scope of heterogeneous

canopies over the floodplains.

Many of the previous researchers adopted numerical simulation techniques in order to clarify the turbulent flow

structures through the vegetation in open channels. For example, Nodaoka and Yagi [23] and Su and Li [24] applied

Large Eddy Simulation (LES) technique for studying the turbulent flows in open channel in the presence of vegetation.

A non-linear k-epsilon (k−ε) model was implemented by Jahra et al. [25] to clarify the mean velocity distributions and

turbulent features. Kang and Choi [26] developed a Reynolds stress model (RSM) to investigate the flow structure

with and without considering the effects of vegetation. The Reynolds averaged Navier Stokes (RANS) technique was

adopted by Anjum and Tanaka [27, 28] in which the turbulent flow features through heterogeneous vegetation

configuration were investigated utilizing CFD code FLUENT. Souliotis and Prinos [29] numerically investigated the

effects of vegetation density on the flow stability and turbulence characteristics with the help of FLUENT. Thus, the

involvement of numerical investigating techniques examines the turbulent characteristics more efficiently.

Zhao et al. [30] and Yan et al. [31] experimentally studied the turbulent flow structure through submerged

vegetation in an open channel; whereas the wake structure in the presence of non-submerged vegetation was

investigated by Yu et al. [32]. Zhao and Huai [33] utilized LES technique to study the influence of discontinuous and

submerged patches of vegetation on turbulence of flow in an open channel. Moreover, they pointed that simulating the

flow structures within the vicinity of vegetation cylinders is difficult to achieve through experimental investigation.

Thus, overcoming this difficulty is one of the advantages of present numerical study.

The present study is focused to numerically investigate the flow behavior in a compound channel with the

floodplains containing vertically layered vegetation. The main objective of the present study is to simulate the 3D flow

properties by investigating the detailed velocity distribution and turbulence characteristics under a varying condition of

submergence of novel kind of layered vegetation over the floodplains. The RSM is implemented for the simulation

purpose. The present study can help in understanding the hydrodynamics of flow through compound channel with

vegetated floodplains, better forest management in case of floods, suitable habitat from ecological point of view, and

three-dimensional (3D) flow phenomena through a complex vegetation array to better elaborate the natural processes

in a riparian environment.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

862

This paper is divided into three major sections: 1. Materials and Methods: in which governing equations of the

numerical model, validation of the model, and modeling setup are presented in detail, 2. Results and Discussion: where

the vertical and lateral profiles distribution, and spatial distributions of the flow and turbulent characteristics are

discussed, 3. Conclusions: At the end, the major outcomes of this study are concluded.

2. Materials and Methods

2.1. Governing Equations

For the steady and incompressible open channel flow, the RANS equations for the continuity and momentum can

be expressed as given in Equations 1 and 2, respectively.

Continuity equation:

𝜕⟨𝑢𝑖⟩

𝜕𝑥𝑖= 0 (1)

Momentum equation:

⟨𝑗⟩𝜕⟨𝑢𝑖⟩

𝜕𝑥𝑗= −

1

𝜌

𝜕⟨⟩

𝜕𝑥𝑖+

𝑣

𝜌

𝜕

𝜕𝑥𝑗(

𝜕⟨𝑢𝑖⟩

𝜕𝑥𝑗+

𝜕⟨𝑢𝑗⟩

𝜕𝑥𝑖) − 𝜌⟨𝑢𝑖′𝑢𝑗′ ⟩ (2)

Where 𝑖 represent the time-averaged velocity in xi direction, 𝑗 represent the time-averaged velocity in xj direction, 𝑣

represent the kinematic viscosity, 𝜌 represent the density of water, represent the pressure, and −𝜌⟨𝑢𝑖′𝑢𝑗′ ⟩ represent

the Reynolds stresses.

The general form of the Reynolds stresses is expressed in Equation 3 [34]. It involves different terms

characterizing the partial differential equation for the independent Reynolds stresses transport.

𝜕𝑅𝑖𝑗

𝜕𝑡= 𝑃𝑖𝑗 + 𝐷𝑖𝑗 − 휀𝑖𝑗 + ∏ + 𝑖𝑗 Ω𝑖𝑗 − 𝐶𝑖𝑗 (3)

Where 𝜕𝑅𝑖𝑗

𝜕𝑡 is the rate of Reynolds stresses, 𝑃𝑖𝑗 is the production rate of Reynolds stresses, 𝐷𝑖𝑗 is the stresses transport

due to diffusion, 휀𝑖𝑗 is the rate of dissipation of stresses, Π𝑖𝑗is the stresses transport due to turbulent pressure strain

interactions, Ω𝑖𝑗 is the stresses transport due to rotation, and 𝐶𝑖𝑗is the convection transport.

The production term and the diffusion terms are modelled as (Equations 4 and 5, respectively):

𝑃𝑖𝑗 = − (𝑅𝑖𝑚𝜕⟨𝑢𝑗⟩

𝜕𝑥𝑚+ 𝑅𝑗𝑚

𝜕⟨𝑢𝑖⟩

𝜕𝑥𝑚) (4)

𝐷𝑖𝑗 =𝜕

𝜕𝑥𝑚(

𝑣𝑡

𝜎𝑘

𝜕𝑅𝑖𝑗

𝜕𝑥𝑚) (5)

Where 𝜎𝑘 = 1.0, and:

𝑣𝑡 = 𝐶µ𝑘2

𝜀 (6)

Where 𝐶µ = 0.09.

Equation 7 shows the modeling of dissipation rate.

휀𝑖𝑗 =2

3휀𝛿𝑖𝑗 (7)

Where 휀 represent the dissipation rate of turbulent kinetic energy and 𝛿𝑖𝑗 represents the Kronecker delta which is

expressed as 𝛿𝑖𝑗 = 1 if i = j and 𝛿𝑖𝑗 = 1 if i ≠ j.

The rotation term is given in Equation 8.

Ω𝑖𝑗 = −2𝜔𝑘(⟨𝑢𝑗′𝑢𝑚′ ⟩𝑒𝑖𝑘𝑚 + ⟨𝑢𝑖′𝑢𝑚′ ⟩𝑒𝑗𝑘𝑚) (8)

Where 𝜔𝑘 represents the rotation vector and 𝑒𝑖𝑗𝑘represents the alternating symbol; 𝑒𝑖𝑗𝑘 = +1 if i, j and k are different

and in cyclic order, e𝑖𝑗𝑘 = −1 if i, j and k are different and in anti-cyclic order; and e𝑖𝑗𝑘 = 0 if any two indices are

the identical.

The pressure strain term is expressed in Equation 9.

∏ = −𝐶1𝜀

𝑘(𝑅𝑖𝑗 −

2

3𝑘𝛿𝑖𝑗) − 𝐶2(𝑃𝑖𝑗 −

2

3𝑃𝛿𝑖𝑗) 𝑖𝑗 (9)

Where 𝐶1 = 1.8 and 𝐶2 = 0.6

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863

Also, the turbulent kinetic energy “k” is modelled in Equation 10.

𝑘 = 1

2 (⟨𝑢𝑖

′2 ⟩ + ⟨𝑢𝑗′2 ⟩ + ⟨𝑢𝑘

′2 ⟩) (10)

And, the convective term is simply modelled in Equation 11.

𝐶𝑖𝑗 =𝜕(𝜌⟨𝑢𝑘⟩⟨𝑢𝑖′𝑢𝑗′ ⟩)

𝜕𝑥𝑘 (11)

2.2. Experimental Setup for Model Validation

The numerical model was validated with the experimental data of Takuya et al. [1]. They conducted a flume study

in a 4.8m long and 0.8m wide channel. The test flume was adjusted to simulate a symmetric trapezoidal compound

cross section with partial vegetation belt on the edges of floodplains. The bed slope was fixed to a value of 1/1000.

The rigid vegetation cylinders were arranged in a staggered pattern with vegetation density (λveg) of 2 m-1. The height

(hv) and diameter (D) of each cylinder were 6 cm and 0.6 cm, respectively. The longitudinal and transverse spacing

between the cylinders was equivalent to ΔS = 5.5 cm. The discharge used in the considered case was Q= 10 L/s

corresponding to the emergent vegetation. The experimental conditions are given in Table 1.

2.3. Boundary Conditions

The modeled geometry for validation was simplified to a reduced length ratio of 1/3 due to small vegetation size

and large mesh. A computational domain of 1.6 m length was modeled, while all the other dimensions were kept the

same as shown in Figure 1(a) (top view) and Figure 1(b) (lateral view). An unstructured mesh with tri-pave scheme

was used which provided 1.4 million grid points. A periodic boundary condition was adopted at the inlet/outlet of the

domain which offered an interface (translational periodicity) between the inlet and outlet of the domain. A symmetry

boundary condition was used at the free surface, and a wall boundary with no-slip condition was applied at the domain

bed, side walls and cylinder walls.

Table 1. Experimental conditions (Takuya et al. 2014 [1]), where hf is the flow depth in floodplain, hm is the flow depth in the

main channel, λveg is the vegetation density, hv is the height of the cylinder, ΔS is the spacing between the vegetation

cylinders, Q is the discharge, U is the average velocity, Fr is Froude number, Re is Reynolds number for cylinder and Re* is

flow Reynolds number.

hf hm λveg D hv ΔS Q U Fr Re Re

*

(cm) (cm) (m-1

)

(cm) (cm) (cm) (L/s) (m/s)

5.75 9.25 2.0 0.6 6 5.5 10 0.181 0.220 1086 12507

Figure 1. Experimental scheme (Takuya et al. 2014 [1]), (a) plan view showing partly covered vegetation on floodplains, and

(b) lateral view

(a)

(b)

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The numerical simulations were performed with a computational fluid dynamics (CFD) code FLUENT. The

turbulence closure was achieved with a 3D RSM. The coupling between pressure and velocity was developed with the

SIMPLE method. The solution was considered to be converged after all the residuals were reached to 1×10-6. The

standard wall function was applied as a near wall treatment. The nodes occurring in cross-wise and depth-wise

directions were doubled to test a mesh independent trial. The variations in the results of primary velocities were less

than 1% due to mesh refinement, which indicated that the mesh independent results are achieved.

2.4. Validation of Numerical Model

The channel cross section was considered axis-symmetric to the central vertical axis. Figures 2(a) and 2(b) shows

the comparison of computational and experimental data of surface and depth averaged velocities along the cross-

section X (see Figure 1a). The vertical axis shows the stream-wise velocity, whereas the horizontal axis shows the

transverse distance in y-direction. It can be observed by both experimental and numerical results that the flow

velocities are significantly reduced in the vegetation part of the floodplain i.e. 0.13 cm < y < 0.25 cm (Figure 2a-b). On

the contrary, the velocities are visibly higher in the non-vegetation part of the floodplain i.e. 0 cm < y < 0.13 cm, as

well as in the main channel i.e. 0.25 cm < y < 0.40 cm. This identifies that the presence of vegetation on the edge of

floodplains can offer noticeable resistance to the flow and can affect the discharge carrying capacity of the main

channel.

The results show that the computational data is in close agreement with that of the experimental data,

demonstrating the validity of the present numerical model. However, the computational results show a minor

difference to that of experimental results in the form of slightly over estimation of the velocity values, which may be

due to the error caused by the RSM simplification.

(a) (b)

Figure 2. Comparison of experimental and computational data of (a) surface velocity and (b) depth averaged velocity along

the half width at cross section X

1.1. Model Setup for Present Study

For the present study, similar dimensions of domain i.e.1.6 m long and 0.8 m wide, were considered as were used

for the validation purpose. The floodplains of the compound channel were consisted of layered vegetation (short

vegetation of 6cm height, and tall vegetation of 12 cm height) of same diameter i.e. D= 0.6 cm and spacing i.e. ΔS=

5.5 cm, which covered full width of the floodplains. The scheme of the computational domain is shown in Figure 3(a).

Two discharge conditions of Q= 10 L/s and Q= 18 L/s were considered in order to simulate the behavior of riparian

(floodplain) vegetation of layered configuration under both emergent and submerged condition of short vegetation

considering low level and high-level flood, respectively.

The lateral views of both cases configuration (Case A and Case B) are shown in Figures 3(b) and (c). All the

boundary conditions for the simulations were kept same as those of validation case. The adopted unstructured mesh

used for the present study gave 2.4 million grid points. The hydraulic and geometric conditions are detailed in Table 2.

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865

(a)

(b)

(c)

Figure 3. Schematic diagram of computational domain, (a) plan view showing layered vegetation on floodplains, and lateral

views for (b) Case A, and (c) Case B

Table 2. Hydraulic conditions for present study, where hvs is the height of short cylinder, hvt is the height of tall cylinder.

Case hf hm λveg D hvs hvt ΔS Q U

Fr Re Re*

(cm) (cm) (m-1)

(cm) (cm) (cm) (cm) (L/s) (m/s)

A 5.75 9.25 2.0 0.6 6 12 5.5 10 0.181 0.220 1086 12507

B 10 13.5 2.0 0.6 6 12 5.5 18 0.202 0.193 1212 22503

For the measurement of flow characteristics, important locations (L1-L5) on a lateral cross section Y were adopted

(see Figure 3). These locations are located in the center region of floodplains (L1 and L2), at the beginning and end of

the floodplains (L4 and L5, respectively), and also in the centerline of the main channel along the width (L3). These

kind of important locations over a cross-section have also been investigated by Yang et al. [18]. Moreover, the flow

structures have also been investigated in the form of spatial distribution of contour plots over the cross-section Y as

well as over the free surface in order to further clarify the flow phenomena.

3. Results and Discussion

3.1. Flow Characteristics

3.1.1. Vertical Profiles Distribution of Velocity

The vertical profiles distribution of mean stream-wise velocity at specified locations for both cases (Case A and B)

is depicted in Figure 4(a-b). The horizontal axis represents the stream-wise velocity (u) which was made non-

dimensional with respect to average velocity (U), whereas the vertical axis represents the depth of floodplain (hf)

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866

which was made non-dimensional with respect to short vegetation height (hvs). It can be observed that the flow

velocities in both cases are significantly reduced within the region of floodplains (L1–L2 and L4-L5); however, there is

no clear difference between the velocities magnitudes measured within the floodplain regions. The velocity

magnitudes are noticeably very high in the main channel region i.e. location L3. This identifies that the vegetation

within the floodplains of compound channel significantly affects the velocity distribution and offers resistance to the

flow. This resistance caused by floodplain vegetation has also been discussed and observed in the previous studies [1,

16]. On the contrary, the velocity in the main stream region without the vegetation increases that could be a

compensation of the resistance to the flow in the floodplains; thus, it can consequently result in an increase in the

discharge carrying capacity of the main channel.

Although, the velocities in both cases close to the bed reduced to minimum due to the resistance offered by the

domain bed; however, difference in the velocity structures exist between Case A and Case B (Figure 4a-b). The

distribution of velocity in Case A remained almost constant above the bed region up-to the height of free surface at all

the locations. This is due to the reason that both the short and tall vegetation were emergent in this case, where an

almost constant drag to the flow was offered by the vegetation structures. This constant distribution of velocity

profiles is identical to that observed by previous researchers of flow investigations through the emergent vegetation

[35, 36]. On the contrary, the simulated velocity profiles in floodplains for Case B shows a different structure due to

the submergence of short vegetation. The velocity distribution above the bed region remained almost constant up-to

the vicinity region of shorter submerged vegetation i.e. hf/hvs≈ 0.8, followed by an inflection point over the top of

submerged vegetation, showing consistency with the results obtained by previous studies [21, 37]. This inflection

point in the velocity profiles over the top of short submerged vegetation is due to the exchange of momentum between

the top of submerged vegetation and the overlying flow. The velocity gradient continued up-to the region just above

the top of short vegetation height i.e. hf/hvs≈ 1.2, and then became constant above it until the free surface. Thus, a

mixing layer over the vicinity region of shorter submerged vegetation (hf/hvs≈ 0.8 to hf/hvs≈ 1.2) is resulted. A mixing

layer is produced when the canopy absorbs sufficient momentum to generate an inflection point in the velocity profile,

which is required for triggering the Kelvin–Helmholtz instability. Thus, the vortices produced due to this instability

dominate the mass and momentum exchange between the canopy and the overlying flow in aquatic canopies [38, 39].

(a) (b)

(c) (d)

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867

(e) (f)

Figure 4. Vertical profiles distribution of: streamwise velocity for (a) Case A, and (b) Case B; transverse velocity for (c)

Case A, and (d) Case B; and vertical velocity for ((e) Case A, and (f) Case B. For the specified locations, see Figure 3

Moreover, the velocity distributions followed almost S-shaped pattern on floodplains when short vegetation was

submerged (Figure 4b). However, an almost logarithmic profile is predominantly observed in the main channel (L3).

These distributions are in agreement to the experimental results of Yang et al. [18]. The velocities in the overlying

flow above the short vegetation height i.e. hf/hvs>1, are increased by a percentage difference of 18-30% due to the

reason that less resistance is offered by the sparse arrangement of tall emergent vegetation in this region, as compared

to the region within the height of short vegetation i.e. hf/hvs< 1.

The vertical profiles distribution for mean transverse velocity (v) and vertical velocity (w) for both cases is

depicted in Figure 4(c-d) and Figure 4(e-f), respectively. The results show that the velocity components in the

transverse and vertical directions are almost minimum i.e. close to zero, at all the locations, in comparison to that of

stream-wise velocity components (Figure 4a-b). However, fluctuations in the velocity profiles are observed at the

locations in the floodplains (L1–L2 and L4-L5) due to the influence of vegetation structures, whereas no fluctuations in

the velocity profile at the location in the main channel (L1) is observed. These fluctuations in the transverse and

vertical velocities have also been observed by previous researchers [22, 27]. The fluctuations in the velocity profiles

are observed to be slightly larger in the transverse direction as compared to those in the vertical directions for both

cases (Case A and B). Moreover, the velocities components in the transverse and vertical directions are considerable

as they indicate the secondary flow and lateral exchange of momentum.

3.1.2. Spatial Distribution of Velocity

The contour plots distribution of mean stream-wise velocity over the cross-section Y is presented in Figure 5(a-b).

It can be noticed that the velocities are reduced to minimum close to bed due to the resistance offered by it. A clear

difference between the floodplains and main channel regions can be observed. The flow in the main channel i.e. 25 cm

< y < 55 cm, experienced larger velocities, where the maximum values of velocity were observed to be in the center

region of the main channel i.e. y ≈ 40 cm. At the interface between the main channel and the floodplain regions i.e. y ≈

25 cm and i.e. y ≈ 55 cm, the flow instability in the lateral direction is triggered by the flow shear due to the presence

of vegetation over the floodplains, which results in the formation of coherent vortices and exchange of momentum [40,

41]. The velocities in the floodplain regions are significantly reduced due to the drag offered by the vegetation. Thus,

large exchange of momentum from the main channel (having high velocity) to the floodplains (having low velocities)

is expected to occur. Moreover, these low velocity regions could significantly encourage the deposition of sediments

in the vegetated floodplains. Within the floodplain regions, a vertical mixing layer is also observed in Case B (Figure

5b) over the region of short submerged vegetation due to the vertical exchange of momentum over this region of flow,

showing consistency with the results observed in Figure 4b. However, the velocity distribution is almost constant in

Case A (Figure 5a) where the vegetation canopy is emergent.

Figure 5(c-d) depicts the spatial distribution of flow velocity over the free surface for both cases. The discharge

passing through the main channel can be easily differentiated from the discharge passing through the vegetated

floodplains. The velocity rapidly increased from the vegetated floodplain regions to the main channel region, followed

by a significant transition and inflection along the interface between these regions. Then, coherent vortices resulted

due to the flow instability dominate the lateral exchange of mass and momentum between the main channel and

vegetated floodplain regions [42, 43]. Moreover, the velocities are reduced to minimum in the regions directly

downstream of the vegetation structures, followed by vortices and wakes i.e. primary Karman vortex streets, which

disappeared after travelling some distance behind the individual vegetation structures. The flow velocities failed to

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868

recover their normal patterns as the influence of vegetation remains for a specific distance behind the downstream

edge. From ecological point of view, aquatic organisms find these regions of low flow velocity suitable for their

physical environment and growth [44, 45]. Sediment deposition also occurs in these regions. Previous researchers also

pointed out that the wake regions behind the vegetation are the regions of fine particle deposition that promote further

growth of the vegetated region [46, 47].

Colourmap of u/U (ms-1/ms-1) for Case A and Case B

(a)

(b)

(c)

(d)

Figure 5. Spatial distribution of mean stream-wise velocity (u/U) over cross-section Y for (a) Case A, and (b) Case B; and over free surface for (c) Case A, and (d) Case B. For the cross-section Y, see Figure 3a

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3.1.3. Lateral Profile Distribution of Velocity

The variations of the depth averaged mean stream-wise velocity along half of the cross-section Y (due to

symmetry) is depicted in Figure 6. The data for both cases further clarifies that the main channel without vegetation

i.e. 25 cm < y < 40 cm, has higher flow velocities compared with the vegetated floodplain region i.e. 0 cm < y < 25

cm, identifying the effect of vegetation and resistance due to it. Within the floodplain region, rise and fall in flow

velocities can be observed, which is due to the influence of the vegetation; where slightly higher velocities are present

in adjacent regions of vegetation structures and comparatively lower velocities occurred in the regions directly

downstream of vegetation, as can be observed in Figure 5. Along the interface (represented by a dashed line), a strong

rise in depth averaged velocity can be observed while moving from floodplain region to the main channel region,

resembling the compensation of higher flow resistance due to vegetation in the floodplain. A remarkable velocity

difference is noticed between the main channel and vegetated floodplain. The discharge (calculated through the

velocity data) passing through main channel of the compound section increased by a percentage difference of 73% and

67% for Case A and Case B, respectively, as compared to the discharge through the floodplains, which favors the

increasing conveyance capacity of compound or natural river channels in severe flood conditions. Moreover, the depth

averaged flow velocities are also observed to be slightly higher i.e. 9% larger, in the floodplain region for Case B, as

compared to that of Case A. This may be due to the reason of higher initial flow velocity or reduction of the drag by

the sparse arrangement of tall emergent vegetation in Case B.

Figure 6. Variations of depth averaged mean velocity (um/U) along the cross-section Y

3.2. Turbulent Characteristics

3.2.1. Vertical Profiles Distribution of Reynolds Stress

Reynolds stresses (that are used to represent turbulence characteristics) including normal stresses (u’u’, v’v’, w’w’)

and shear stress (-u’w’) measured at specified locations are presented in Figure 7(a-h) for both cases (Case A and B).

All the stresses were made dimension-less with respect to U2. In Figure 7(a-h), u’ indicates the velocity fluctuations in

streamwise direction, v’ indicates the velocity fluctuations in transverse direction, and w’ indicates the velocity

fluctuations in vertical directions. Larger values of Reynolds stresses accumulated close to the bed region for both

cases, showing consistency with those observed in the previous research work by Anjum et al. [22]. Whereas, while

moving above the bed region, the Reynolds stress distribution became uniform at almost all the locations.

(a) (b)

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870

(c) (d)

(e) (f)

(g) (h)

Figure 7. Vertical profiles distribution of Reynolds normal stresses (u’u’, v’v’, and w’w’) for (a,c,e) Case A, (b,d,f) Case B,

and Reynolds shear stress (-u’w’) for (g) Case A, and (h) Case B

Moreover; in case B, a slight modulation in the Reynolds stress profiles around the mixing layer region i.e. 0.8 <

hf/hvs < 1.2, is observed for the locations in floodplain regions (L1-L2 and L4-L5). However, the distribution of stresses

again became constant above the mixing layer region i.e. hf/hvs > 1.2. The results also show slightly less accumulation

of Reynolds shear stress close to the bed region (Figure 7g-h), as compared to other Reynolds normal stresses (Figure

7a-f). In addition to this, the bed region acquires high Reynolds stresses in the floodplain regions (L1-L2 and L4-L5), as

compared to the main channel region (L3). The sharp spatial variation in the Reynolds stress influences the dynamics

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871

of sediments, which affects the particle size distribution of the sediments available in the flow. Thus, these regions of

influenced Reynolds stresses near the bed could encourage the deposition of sediments in the canopy regions [33].

3.2.2. Lateral Profile Distribution of Turbulent Kinetic Energy

The variation in the dimensionless depth-averaged Turbulent Kinetic Energy (TKE, Figure 8) along the cross-

section Y further depicts the difference of turbulent flow structure between the vegetated floodplain and main channel

without the vegetation. Within the vegetation region of floodplains, fluctuations in the depth-averaged TKE can be

observed, followed by a saw-tooth dissemination, which showed consistency with the previous research works [33,

48]. Along the interface of floodplain and main channel i.e. y ≈ 25 cm, for both cases (Case A and B), a peak in the

TKE is observed due to the interaction of slow and fast flow over this region which caused a larger turbulence. On the

contrary, the distribution of TKE is observed to be uniform in the main channel region. This identifies the existence of

turbulence more on the vegetated region of floodplains as compared to the main channel region. The production of depth-averaged TKE is also found to be slightly higher for Case A due to the larger hindrance offered by the overall

emergent canopy, as compared to that in Case B.

Figure 8. Variations of depth averaged turbulent kinetic energy (TKE)along the cross-section Y.

3.2.3. Vertical Profiles Distribution of Turbulent Intensity

Figure 9 (a-b) shows the turbulent intensity (%) profiles at the specified locations for all the cases. It can be

witnessed that the percentage of turbulence is highest close to the bed region at almost all the locations for Case A and

Case B. Along the depth of flow above the bed region, the intensity percentage decreased logarithmically at location

L3 as no resistance due to the vegetation was observed in the main channel. On the contrary, the vertical distribution of

turbulent intensity at other locations (L1–L2 and L4-L5) remained almost uniform along the flow depth within the height

of short vegetation i.e. hf /hvs < 1, for both cases (Case A and B). Moreover, a slight inflection in the intensity profiles for the locations in the floodplains (L1–L2 and L4-L5) is observed around the top of shorter submerged vegetation i.e. hf

/hvs ≈ 1, for Case B (Figure 9b), which is due to the variation of vegetation density and drag offered by it. This shows

consistency with those profiles of velocity in Figure 4b. Furthermore, the turbulent intensity at locations in the

floodplains (L1–L2 and L4-L5) is more influenced by vertically layered vegetation as compared to the location L3 in the

main channel.

(a) (b)

Figure 9. Vertical profiles distribution of turbulent intensity (%) for (a) Case A, and (b) Case B

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3.2.4. Spatial Distribution of Turbulent Intensity

Spatial distribution of turbulent intensity is reported here to study the turbulence in the compound channel with

layered vegetated floodplains. Figure 10(a-b) shows the contour plots of turbulent intensity for all the cases (A–B)

along the cross-section Y. The spatial variation of the intensity distribution is high in the vegetated floodplains,

indicating that the flow structure is non-uniform in these regions. Thus, larger percentage of turbulence is observed in

the floodplain regions, which is due to the presence of vegetation. Within the region of above the short vegetation

height for Case B (Figure 10b), the percentage of turbulence significantly decreased due to the reduction in drag by the

vegetation. A clear difference is observed between the floodplain regions and the main channel region. Close to the

bed region of main channel, the amount of percentage is also observed to sufficiently increase due to the resistance

offered by the bed, showing correspondence to the low velocity observed in Figure 5(a-b). Moreover, the intensity in

the main channel tries to achieve a uniform distribution above the bed region. Along the interface of the main channel

and floodplains, a strong turbulence in the flow can be observed which is due to the reason of flow instability and

exchange of momentum over this region of flow.

To further clarify the turbulent flow structure in the compound channel, contour plots of intensity over the free

surface have been plotted and presented here (Figure 10c-d). It further shows the influence of layered vegetated

floodplains on the flow turbulence. Higher percentage of turbulent intensity is observed directly behind the vegetation

cylinders in both cases (Case A and B) due to high flow resistance in these regions, followed by trailing vortices

which requires some sufficient distance to reach in the stable state. A stronger rise in the flow turbulence over the

interface region of floodplains and main channel is found, followed by a decrement in the turbulence percentage while

moving away from the floodplains towards the main channel. This effect on the interface region of floodplains and

main channel has also been studied in the previous researches [16, 18]. This signifies the promotion of lateral

exchange of momentum at the boundary of main channel and floodplains. This kind of complex flow structure around

the vegetation vicinity and interface of both regions is difficult to capture in an experimental study [33, 49]. Thus, it

demonstrates the significant advantage of present numerical study. Moreover, due to high initial flow velocity and

Reynolds number, the overall turbulent intensity in Case B is found to be higher i.e. 20%, as compared to that of Case A.

Colourmap of T.I (%) for Case A

(a)

Colourmap of T.I (%) for Case B

(b)

(c)

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873

(d)

Figure 10. Spatial distribution of turbulent intensity (%) over cross-section Y for (a) Case A, and (b) Case B; and over free

surface for (c) Case A, and (d) Case B

4. Conclusions

The present numerical study investigated the impact of vegetation on the compound channel to clarify the velocity

distribution and turbulence characteristics over the floodplains with layered vegetation configuration. To achieve the

desired objectives and computing the flow structures through vegetated compound channel, a CFD code FLUENT was

used for simulation and RSM model was used for the turbulence closure. The following conclusions were drawn from

the current study:

When the short vegetation became submerged during high flows, an inflection point in the velocity distribution

occurred resulting in a significant mixing layer over the top of submerged vegetation due to the vertical exchange of momentum between the top of submerged canopy and the overlying flow. Whereas, when both the

short and tall vegetation remained emergent during low flows, the mean flow characteristics showed almost

uniform distribution along the depth of flow.

The flow velocities significantly reduced in the floodplain region due to the drag offered by the layered

vegetation. On the contrary, the velocities as well as the discharge carrying capacity of the main channel

increased by a percentage difference of 67-73%. A flow separation followed by a lateral exchange of momentum

resulted along the interfacial region of vegetated floodplains and the main channel.

The Reynolds stress distribution and turbulent intensity showed almost uniform distribution above the bed region

up-to the vicinity of top of short vegetation height, followed by a slight inflection point over the mixing layer

region. Moreover, a uniform distribution of Reynolds stress and low accumulation of turbulent intensity in the

floodplain regions indicated a positive response for the sediment deposition as well as for the nourishment of the aquatic organisms in the riparian environment.

The peak values of turbulent kinetic energy and turbulent intensity occurred over the interfacial region of

floodplains and the main channel, signifying a strong turbulence over this region due to high instability of the

flow that could promote lateral exchange of momentum over the boundary.

The present numerical study successfully replicated 3D flow behavior through a vegetated compound channel. It

can help in understanding the resistance offered by the vegetation over the floodplains and give a clear understanding

of the discharge carrying capacity of the compound channels. The flow hydrodynamics explored through this research

could be implemented to natural riparian environment. The knowledge of resistance due to vegetation drag can help in

designing effective measures to reduce the shear force that acts on bed sediment particle, structural analysis and

depreciation on the land characteristics. Furthermore, the vegetation resistance over interfacial region will also

significantly affect the momentum exchange and it can help in preventing inundation if applied wisely. Thus, a proper

management and wise use of the vegetation would provide us a new strategy of sustainable flood protection. In

addition, this study can be used to enhance our knowledge of vegetation patterns for future planning of flood

protection measures and the outcomes of this study may become useful while designing ecological habitats.

More study to investigate the effect of vegetation density and varying heights of vegetation over the floodplain is

required to further clarify the phenomena. In the future, the flow and turbulent characteristics for several other patterns

of rigid vegetation on floodplains are required to be studied to better understand the natural environment such as

natural streams and rivers.

5. Conflicts of Interest

The authors declare no conflict of interest.

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3. References

[1] Takuya, U., Keiichi, K., and Kohji, M. "Experimental and numerical study on hydrodynamics of riparian vegetation". J.

Hydrodyn, 26, (October 2014): 796-806. doi: 10.1016/S1001-6058(14)60088-3.

[2] Bousmar, D., and Zech, Y. "Momentum transfer for practical flow computation in compound channels". J. Hydraul. Eng, 125,

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Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

877

The Effects of Nano Bentonite and Fatty Arbocel on Improving

the Behavior of Warm Mixture Asphalt against Moisture

Damage and Rutting

Sepehr Saedi a*

, Seref Oruc b

a Assistant Professor, Department of Civil Engineering Avrasiya University, Trabzon, Turkey.

b Professor, Department of Civil Engineering Karadeniz Technical University, Trabzon, Turkey.

Received 05 February 2020; Accepted 04 April 2020

Abstract

The use of warm mix asphalt (WMA) technology has increased dramatically in recent years to protect the environment

and reduce energy consumption. Despite numerous advantages, WMAs are less commonly used as a result of their lower

performance in comparison to HMAs. One of the main reasons for the low performance of WMAs is their high moisture

sensitivity. In recent decades, bitumen modifiers have been used to improve the performance of asphalt mixtures. One of

the additives that has recently been used to modify the characteristics of bitumen, is bentonite. The grade of asphalt

cement used in this study is PG 64 -22 and the Bitumen is modified with 1, 3, 5 and 7% nano bentonite. Also, 0.3% fatty

Arbocel has been used for the preparation of WMA. Indirect tensile strength (ITS) test and Nicholson stripping test are

used to determine moisture sensitivity and dynamic creep test and LCPC are also used to evaluate the rutting potential.

The results indicate that, increasing the percentage of nano bentonite and applying 0.3% of fatty Arbocel improves the

resistance of mixture against moisture damage. Also it was found that increasing the mixture hardness decreases the

permanent displacement and rutting potential of WMAs. So, it is suggested that the consumption of these additives

increases WMA’s lifetime and decreases its maintenance cost.

Keywords: Warm Mix Asphalt; Dynamic Creep Modules; Rutting; ITS.

1. Introduction

Possible failures in WMAs are divided into four major groups: 1. Permanent deformation or rutting, 2. Fatigue or

load associated cracking, 3. Low temperature or thermal cracking, 4. Moisture damages. WMA is an emerging

technology, In line with concerns about global warming and energy consumption in asphalt industry. In WMA

production, the mixing and compaction temperature are reduced from 10 to 38° C

compared to the HMA (Hot mix

asphalt) that is produced and compacted at temperatures of 145 to 150° C

[1]. Production of temperature reduction

results in reduced fuel consumption and manufacturing costs. It also reduces the amount of greenhouse gas emissions

in asphalt production process [2]. Modified bitumen is one solution for improving the pavement performance [3]. Clay

modified bitumen had been used since 100 years ago. As soon as the clay particle-page reaches inside the bitumen, it

creates similar rebar properties inside reinforced concrete which improve bitumen performance against cracking and

deformation in high temperatures [4]. Research shows that adding clay to bitumen increases bitumen softening point

and decreases its elasticity which increases the resistance of hot mix asphalt against thermal cracking [5]. Adding nano

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091514

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

878

clay and nano lime to the hot asphalt mixture increases the durability of the mixture against freeze -thaw cycles [6].

The use of nano-clay also increases surface free energy and fatigue resistance of HMAs [7]. Adding cement and nano

clay together have been found to improve the initial strength and cracking resistance, and reduce moisture

susceptibility of cold mix asphalts [8]. Bentonite is a type of clay with high plastic and colloidal properties; it mainly

consists of montmorillonite minerals [9]. Studies on HMA containing nanoclay modified bitumen show that nanoclay

improves the performance of HMA against moisture damages and rutting [10]. Bentonite improves the rheologic

characteristics of asphalt cement and it against aging [11].

Researches on HMAs containing bentonite modified bitumen indicate that these types of mixtures have higher

shear strength and longer fatigue life than conventional HMAs [12]. Bentonite improves the rheological properties of

bitumen at low temperatures and increases the resistance of hot asphalt mixtures against thermal cracks [13]. Semi-hot

additives such as sasubit and asphamine have no significant effect on the amount indirect tensile strength and moisture

sensibility of WMAs [14]. Nano-Zycotherm improves the moisture sensitivity of WMAs, but its effect on failure of

rutting has’nt been reported. Sasubit, are found to increase the hardness and modulus of WMAs and decreases their

permanent displacements [15]. Recent studies have proved that the addition of Evotherm and paraffin wax was not

effective against moisture susceptibility [16]. Also, based on research findings, the use of rubber enhanced resistance

against permanent deformation of WMAs but could not affect fatigue resistance [17]. Warm mix asphalts containing

styrene–butadiene–styrene (SBS) copolymer were found to have a good behavior against rutting damage [18]. A

considerable number of researches have been done on nanoclay modified polymers. Variables that have a great impact

on the final nanocomposite include the choice and the type of used clay, the components of the polymer used, and how

to mix them [19]. Natural and synthetic fibers can improve the resistance of fracture and the performance of WMAs

[20]. Studies show that cellulose fibers improve the moisture sensitivity of hot asphalt mixtures and help the adhesion

between aggregates and bitumen [21]. Fibers containing Arbesol improves SMA resistance against failure and

deformation and increases the efficiency of pavements [22].The additives containing Arbesol ZZ 8/1 take less energy

for compaction of SMAs compared to nonadditive mixtures and improves its resistance against rutting, fatigue, and

moisture sensitivity [23]. Synthetic macrofibres can help WMA’s rutting resistance and improve their performance

and life spans [24]. Based on the above-mentioned studies, the current study is aimed at performing a laboratory

investigation on WMA mixtures containing nanobentonite and fibers through a mechanistic–empirical approach to

determine the effect of different additives on increasing the service life of pavements.

Figure 1. Research methodology

The Work Program

Aggregate

Mechanical Tests Asphalt Cement

Tests

Additive

Mechanical Tests

Warm Mix

Asphalt Provide

Marshall Test Moisture Test Rutting Test

1. Marshall Stability

2. Flow

3. Marshall

Quotient

1. Nicholson Stripping Test

2. Moisture Damage Test

1. Dynamic Creep Test

2. LCPC Test

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2. Materials and Methods

2.1. Applied Material

The aggregates used in the current study were extracted from a limestone obtained from a mine in the southern part

of Trabzon province in Turkey. Filler was used to provide WMA mixtures from stone powder, which has been found

to reduce moisture sensitivity in previous studies. Mechanical tests were adminstered to determine the quality of

aggregates according to the common standards which are presented in Table 1.

Table 1. Mechanical properties of aggregates

Properties Method Requirement Values

Water absorption % ASTM C127 [25]

2.8 Max. 0.48

Los Angeles abrasion (%) ASTM C131 [26] 30 Max. 18

Flat and Elongated ASTM D 4791 [27] 20 Max. 12

Coarse aggregate specific density (g/cm3) ASTM C127 [25] - 2.721

Fine aggregate specific density (g/cm3) ASTM C127 [25] - 2.731

Mineral filer specific density (g/cm3) ASTM C127 [25] - 2.741

The type of Asphalt cement in the current study was PG 64 -22, the properties of which are described in Table 2.

Table 2. Physical properties of PG 64 -22

Properties Method Requirement Values

Penetration (0.1 mm) EN-1426 [28] - 58

Softening point (R&B) (°C) EN-1427 [29] - 50.2

Table 3 presents the physical properties of bentonites used in the current study.

Table 3. Physical properties of Bentonite

Gs LL PL PI

2.233 165 43 122

Table 4 illustrates the Fatty Arbocel used in the current stud.

Table 4. Properties of Fatty Arbocel

Properties Values

Fiber length (mm) 2> and <5

Fiber thickness (mm) 2

Bulk density (g/cm3) 410-440

Flash point (°C) >300

2.2. Samples Provided

To determine optimum bitumen content according to the Marshall method, three samples were prepared for each

asphalt concrete containing 4, 5.4, 5, 6 and 6.5% bitumen. To make WMAs, the aggregates were heated at 135° C

for

24 hours. Bitumen was heated up to 130° C

to determine the optimum bitumen percentage mixture. Furthuremore,

Marshall hammer was performed on both sides of compacted Specimens of 75 impacts to stimulate heavy traffic. The

optimum amount of bitumen content (OBC) for the WMAs was also determined 5.8 %. Next, the amounts of bentonite

additives which were 1, 3, 5, 7 % of the total weghit of OBC were added to the mixture, and finally the fatty Arbocel

fiber which was 0.3 % of total sample weight was combined.

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880

Figure 2. Marshall Specimens

2.3. Marshall Test

The Marshall quotient indicates the rigidity of the asphalt mixture. Any increase in the ratio i, increases the rigidity

and resistance of the mixture against permanent deformations. Thus, to estimate the ratio, the specimens are placed

inside Marshall Jacket and pressurized to evaluate Marshall Stability and flow values [30].

Figure 3. Marshall Test

2.4. Moisture Damage Test

According to AASHTO T283-03 standard, a number of six samples were made, out of which three were used for

indirect tensile strength test under dry conditions (unsaturated) and the remaining three samples were used for testing

under saturated conditions.

ITS =2Pmax

πDt (1)

Where;

ITS: the tensile strength (kpa), P: maximum load (N), t: specimen thickness (mm), D: the specimen diameter (mm).

TSR % = ITSwet

ITS dry× 10

(2)

As mentioned above, the minimum values of TSR for resistance against water damages should be equal to 75% [31].

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Figure 4. Moisture Damage Test

2.5. Stripping Test

Nicholson stripping test was carried out to determine the resistance of bitumen against separation of aggregates due

to the water effect. The adhesion between the aggregate and asphalt cement plays an important role in increasing the

resistance of asphalt mixture against permanent deformation which consequently increases its durability [32].

Figure 5. Nicholson Stripping Test

2.6. Dynamic Creep Test

It is an experiment to measure permanent deformations hat occur in asphalt concrete under the influence of

repeated loads. During dynamic creep test, the specimens are subjected to a uniaxial compressive load at a specified

period. After repeating each load, permanent deformations in the samples are measured [33].

Ɛc = 3n − L1)/G (3)

Ɛr = (L2 − L3 ) /[G − (L3 − L1)] (4)

= F/A (5)

Ec = / Ɛc (6)

Er = / Ɛr (7)

Where;

Ɛc: Plastic deformation, Ɛr: Elastic deformation, Ec: Creep module, Er: Elastic module, : Stress (kPa)

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Figure 6. Dynamic Creep Test

2.7. LCPC Rutting Test

In order to perform a French rutting test, slab-shaped specimens with a length of 500 mm and a width of 180 mm

and a height of 50 mm are prepared for each WMA mixture. The measurements are performed at 1000, 3000, 5000,

10,000, 30,000 and 50,000 rpm. Finally, using equation, the amount of rutting settlement is determined for each

mixture (8), [34].

Y = A[N/1000]B (8)

Where;

Y: N cycle settlement (mm), A: settlement at 1000 rpm, B: It is the slope of the linear line in logarithmic

coordinates.

Figure 7. LCPC Rutting Test

3. Results and Discussion

3.1. Marshall Test

The results of Marshall Quotient test are given in Figure 8, which presents the improvement of samples’ Marshall

Quotient with the addition of nano-bentonite. This was due to the improvement in the properties of bitumen modified

with nano-bentonite as a result of high specific surface area of nano-particles. The viscosity and adhesion of modified

bitumen have been increased by increasing the amount of nano-bentonite. Furthermore, fatty arbocel helps the

adehesive property between bitumen and aggregates. So, it can be said that using the nano-bentonite and fatty arbocel

can improve the resistance of WMAs against permanent deformations. The results are in line with the results of the

Iskender E. experiments [10]. The results show that Marshall Quotient increased by 2.79, 8.7, 11.4 and 15.6 %

compared to conventional mixture.

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883

Figure 8. Marshall Quotient test results

3.2. Nicholson Stripping Test

The results of Nicholson stripping tests are given in Figure 9.

Figure 9. Results of Nicolson stripping test

Figure 9 reveals that adding 0.3 % of fatty Arbocel to the mixture and increasing the amount of nano-bentonite (1,

3, 5 and 7) increases the resistance of WMAs in comparison to none additive WMAs which showed 33%, 40%, 67%,

75% increase. It is clear from the data that the additives significantly affected the adhesion property of asphalt cement

and improved its resistance against stripping. Hence, it is suggested that WMAs modified with nano-bentonite and

fatty Arbecol are very useful in the performance of them against stripping, especially in cold and wet regions.

3.3. Moisture Damage Test

The Indirect tensile strength (ITS) values of WMAs are presented in Figure 10. As shown, nano-bentonite together

with Arbocel increases the tensile strength of the saturated and unsaturated samples.

518.5 532.8

564.3 577.9

599.4

250

350

450

550

650

WMA01 WMA02 WMA03 WMA04 WMA05

Ma

rsh

all

Qu

eti

en

t (k

g/m

m)

60

45 40 38

34

0

20

40

60

80

100

WMA 01 WMA 02 WMA 03 WMA 04 WMA 05

Nic

hols

on

Str

ipp

ing

(%)

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Figure 10. Results of ITS test

It is observed that the ITS value of the saturated mixtures are lower than that of dry mixtures, due to the presence of

water in the mixture resulting in reduced adhesion between the aggregates and bitumen leading to a reduction in the

strength of the asphalt mixture samples under the loads . Nano-bentonite improves the elastic properties of bitumen and

polymer and Arbocel improves the elastic properties of the mixture. Therefore, the combined application of these

additives can play an important role in increasing the tensile strength of WMAs and compensate for the weakness of

these mixtures against moisture damages.

The TSR results in Figure 11 show that the samples containing the additives were able to pass the standard criteria,

wheras the non-additive samples remained below the criteria specified for the TSR. Anti-stripping properties of nano-

bentonite and Arbocel increase the TSR of modified mixtures. According to the results presented in Figure 11, the

highest amount of TSR occurs in samples containing 0.7% nano-bentonite and 0.3 % fatty Arbocel, indicating that as

the percentage of nano-bentonite increases, the adhesion to the semi-hot asphalt mixtures also increases. Anti-stripping

properties of additives causes the mixture to resist higher moisture compared to other samples during frost and thaw

cycles. Thus, it is apparent that the additive increases the resistance of the mixtures against moisture damage. These

results are in line with the results of Ameri M. experiments [35].

Figure 11. Results of TSR

5.62 5.86

6.09 6.41

6.78

4.02 4.48

4.74 5.11

5.58

0

1

2

3

4

5

6

7

8

WMA01 WMA02 WMA03 WMA04 WMA05

ITS

(K

pa)

ITS dry

ITS wet

71.53 76.45 77.83 79.71 82.3

0

10

20

30

40

50

60

70

80

90

100

WMA01 WMA02 WMA03 WMA04 WMA05

TS

R (

%)

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3.4. Dynamic Creep Test

Dynamic creep test was applied to measure WMAs deformations the results of which are shown in Figure 12.

Figure 12. Creep Module Diagram

According to the results, there was a decrease in WMAs with nano-bentonite and Fatty Arbocel compared to the

mixtures without additives. Mixtures containing nano-bentonite Arbocel exhibit lower creep modulus in the same cycle

than the control mixture, indicating greater plastic deformation in the control mixture. It is suggested that the addition of

nano-clay with arbesol strengthen WMAs against deformation compared to conventional WMAs. As a result, Nano-

bentonite and arbocel with improved adhesion between aggregate and bitumen can improve the creep performance of

WMAs and rutting potential. The results also suggest that, in mixtures containing nano-benonite and Arbocel, the

hardness increased and these mixtures showed higher resistance against rutting compared to non-additive WMAs.

Hence the modified mixture is claimed to have more service life than none modified mixture.

3.5. LCPC Rutting Test

LCPC rutting testing was used to measure the depths of the rut. As shown in Figure 13, the addition of arbesol and

nano-bentonite decreases the depth of the rut.

Figure 13. Results of LCPC Rutting test

0

10

20

30

40

50

60

70

0 10000 20000 30000 40000 50000 60000 70000

Creep

Sti

ffn

es

(MP

a)

Cycle

WMA 01

WMA 02

WMA 03

WMA 04

WMA 05

2

3

4

5

6

7

8

9

10

0 10000 20000 30000 40000 50000

Ru

ttin

g (

%)

Cycle

WMA01

WMA02

WMA03

WMA04

WMA 05

Civil Engineering Journal Vol. 6, No. 5, May, 2020

886

Rutting test was performed on slab form samples. According to the proposed standard for LCPC method, the

amount of rutting caused by 30,000 wheel cycles, is expected to be not bigger than 6%. According to the curve in

Figure 13, it is apparent that samples containing 1, 3 and 5% nano bentonite did not reach the standard limit. By

increasing the amount of nano bentonite, not only we could reach the reach the standard limit, but also we were

successful in reducing the rutting percentage of samples containing 7% nano-bentonite and 0.3% fatty Arbocel about

one second compared to non-additive samples. These results are in line with the results of the Ziari H. experiments

[36]. Given the comparative rutting curves, it is clear that the application of fatty Arbocel fibers and nano-bentonite

reduces the rutting time compared to non-additive samples. According to the curve, samples containing 7% nano-

bentonite and 0.3% fatty Arbocel represented the best results which can yield important achievements in roads with

heavy vehicle traffic.

4. Conclusions

Usage nano-bentonite and fatty Arbecol as additives to WMAs has a significant effect on the Marshall Quotient

of these mixtures compared to non-modified mixtures. The results show that Marshall Quotient was increased by

2.79, 8.7, 11.4 and 15.6 % compared to conventional mixtures. So, it is suggested that using the nano-bentonite and

fatty arbocel can improve WMAs resistance against permanent deformations.

The values obtained from Nicholson stripping tests revealed that nano-bentonite and fatty Arbecol additives were

effective on the adhesion between aggregates and bitumen. Hence, it is suggested that WMAs modified with nano-

bentonite and fatty Arbecol are very useful in their performance against stripping, especially in cold and wet

regions.

Nano-bentonite and Fatty Arbocel increase the resistance of WMAs against moisture damages, According to the

results, the highest amount of TSR occurs in samples containing 0.7% nano-bentonite and 0.3 % fatty Arbocel.

Due to its lower cooking temperature in the preparation process which consequently increases the mixture’s

moisture sensitivity. The use of nano-bentonite and Fatty Arbocel additives can have a very effective role against

the functional weakness caused by moisture sensitivity.

According to the results of the dynamic creep test, both the resistance of modified WMAs against permanent

deformation and the amount of elastic deformation revealed an increase. The highest creep modulus value was

obtained by mixing 7% nano-bentonite and 0.3% fatty Arbecol. The results also revealed that in mixtures

containing nano-benonite and Arbocel, mixtures hardness increased. Also, these mixtures had higher resistance

against rutting compared to non-additive WMAs.

According to the results of the LCPC rutting test, application of fatty Arbocel fibers and nano-bentonite reduces

the rutting process when compared to non-additive samples. Increasing the amount of nano bentonite to the reach

standard limit, the rutting percentage of samples containing 7% nano-bentonite and 0.3% fatty Arbocel decreased

about one second compared to non-additive samples. It is clear that nano-bentonite and fatty Arbecol can help the

performance of WMAs and improve their life span.

5. Conflicts of Interest

The authors declare no conflict of interest.

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Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

889

Post-Fire Behavior of Post-Tensioned Segmental Concrete

Beams under Monotonic Static Loading

Nazar Oukaili a, Amer F. Izzet

a, Haider M. Hekmet

b*

a Civil Engineering Department, College of Engineering, University of Baghdad, Iraq.

b Civil Engineering Department, Al-Farabi University College, Iraq.

Received 31 October 2019; Accepted 09 February 2020

Abstract

This paper presents a study to investigate the behavior of post-tensioned segmental concrete beams that exposed to high-

temperature. The experimental program included fabricating and testing twelve simply supported beams that divided into

three groups depending on the number of precasting concrete segments. All specimens were prepared with an identical

length of 3150 mm and differed in the number of the incorporated segments of the beam (9, 7, or 5 segments). To

simulate the genuine fire disasters, nine out of twelve beams were exposed to a high-temperature flame for one hour.

Based on the standard fire curve (ASTM – E119), the temperatures of 300C (572F), 500C (932 F), and 700C (1292F)

were adopted. Consequently, the beams that exposed to be cool gradually under the ambient laboratory condition, after

that, the beams were loaded till failure to investigate the influence of the heating temperature on the performance during

the serviceability and the failure stage. It was observed that, as the temperature increased in the internal layers of

concrete, the camber of tested beams increased significantly and attained its peak value at the end of the time interval of

the stabilization of the heating temperature. This can be attributed to the extra time that was consumed for the heat

energy to migrate across the cross-section and to travel along the span of the beam and deteriorate the texture of the

concrete causing microcracking with a larger surface area. Experimental findings showed that the load-carrying capacity

of the test specimen, with the same number of incorporated concrete segments, was significantly decreased as the heating

temperature increased during the fire event.

Keywords: Segmental Beam; Post-tensioning; Fire Test; Gradual Cooling; Serviceability; Load Capacity.

1. Introduction

Post-tensioned segmental concrete girders have a significant implementation in bridge engineering due to the

facilities that offered during the construction process. This method of construction has many advantages such as

substantial economical savings due to the possibility of weather-independent segment production and a shorter

construction period, simple element assembly at the job site, replacement ability of deteriorated tendons, the

concreting and prestressing operations are independent, small light segments, profiling of the main external steel is

easier to check, and the friction may be reduced [1]. It is well known that the strength of reinforced concrete and

prestressed concrete members decreases after the exposure to a fire disaster. The main fire safety objectives are to

protect life and prevent failure. Following a fire, if no collapse happens, there is a possibility of fire-induced damage.

It should be noted that the study of the heating history of concrete is very significant to define whether the concrete

structure exposed to fire and its components remain intact from the structural aspect. The evaluation of concrete

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091515

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

890

structures for fire damage commonly starts with the visual inspection of cracking, discoloration, spalling of concrete,

and consequently, determining the residual strength of concrete. The performance of ordinary reinforced concrete and

prestressed concrete members exposed to the fire attack was studied by many researchers [2-8]. Chan et al. [2] studied

experimentally the pore structure and their distribution in addition to the mechanical properties of the normal-strength

(NSC) and high-performance concrete (HPC) after the exposure to a high temperature. The residual compressive

strength of the concrete specimens was examined after subjecting concrete to a temperature of 800 °C. It was

concluded in this study that after exposure to high temperature, the degradation in strength of HPC was more severe

than in NSC.

Phan and Carino [9] presented the data for behavioral variances between the normal-strength concrete (NSC) and

the high-strength concrete (HSC) at high temperatures. Also, they reported the material behavior after fire exposure,

the code provisions for members that subjected to fire attack, and the analytical modeling of HSC at high temperature.

Expanded vermiculite is a significant lightweight aggregate for cementitious materials that are used for fire resistance

applications. Koksal et al. [3] examined four different compound mixtures under high temperatures of 300, 600, 900,

and 1100 °C for 6 hours in varying amounts of expanded vermiculite. They studied the mechanical and physical

properties of concrete including its unit weight, porosity, water absorption, the residual compressive and splitting

tensile strengths, and the ultrasonic pulse velocity after fire exposure and consequently air cooling. Zhang et al. [10]

demonstrated that the mechanical characteristics of prestressing steel during high temperatures and after cooling are

requested for the assessment of the resistance to fire and the residual post-fire load-carrying capacity for the

prestressed concrete structures. In that study, mechanical properties of prestressing wires through and after the

exposure to the fire were examined by sophisticated equipment to upgrade the reliability and accuracy of the material

characteristics database at high temperatures of performance-based design objectives. Accordingly, empirical formulas

were proposed for residual strength. Myers and Bailey [5] examined the residual characteristics of uncoated seven-

wire, 12.7 mm and 9.5 mm diameters low-relaxation grade 270 (1862 MPa) prestressing strands under the excessive

temperature of 260, 427, 538, 649, and 704 °C. The results of that research indicated that there was a loss of tensile

strength of the strand could attain 26.0% when the temperature rose from 538 to 649 °C.

Abdelrahman et al. [6] tested a series of statically determinate prestressed concrete beams under fire to investigate

the effect of different parameters on the behavior of such structural members including the prestressed index, the

concrete compressive strength, and the thickness of the concrete cover. Five specimens were tested in lab conditions

while seven beams were loaded up to its working load and exposed to fire for three hours at 600 °C and left to cool

gradually at the ambient temperature then tested up to failure. Izzet and Al-Dulffy [8] investigated the effect of fire

flame and the rate of loading on the service behavior of partially prestressed concrete beams and the residual strength.

Seven pretensioned concrete beams have been fabricated and tested. One beam was considered as a reference beam

that tested under static loading without fire exposure. Meanwhile, the other six were exposed initially to fire test and

consequently to monotonic static loading, where each pair were subjected to the same heating temperatures of 300,

500 or 700 ᴼC but cooled in different scenario (i.e., gradually or suddenly). It was concluded that the sudden cooling in

comparison to the gradual cooling had a worse impact on the residual load-carrying capacity of the tested specimens.

To date, there are several experimental programs that were carried out to investigate the behavior of segmental

concrete beams under external load. Sivaleepunth et al. [11] and Nguyen et al. [12] presented the results of the

nonlinear finite element analysis on the segmental concrete beams with external tendons. Algorafi et al. [13]

investigated the structural behavior of dry joined externally prestressed segmental beams (EPS) under combined

stresses (bending, shear and torsional stresses). It was found that the reduction in the load-carrying capacity of such

beams can be compensated by the implementation of a higher prestressing force as well as by increasing the stirrup

reinforcement area in the joint regions. To ensure serviceability, the joints in the support regions should still be

sufficiently prestressed under service load conditions. Yuan et al. [14] studied experimentally the performance of

segmental concrete box beams with hybrid tendons. Three scaled-down specimens with different ratios of the number

of internal tendons to the number of external tendons were tested up to failure. Test results showed that as more

internal tendons were used, higher load-carrying capacity and better ductility were achieved. Therefore, the ratio of

these hybrid tendons not less than 1:1 was recommended.

Moubarak et al. [15] investigated the second-order effect in 25 externally prestressed monolithic and segmental

concrete beams. The main parameters of this study were the shear span to depth ratio, the effective prestressing level,

and the profile of external strands. A suggestion for a solution that improves the system efficiency of the segmental

beam with external strands was presented. Thorough analysis based on the mechanics of load transfer, fracture and

damage mechanics were proposed. Jiang et al. [16] conducted a series of tests to investigate the effect of using hybrid

tendons, load location, and the number of joints on the flexural behavior of post-tensioned segmental concrete beams

(PSC) with dry joints. It was noticed that the flexural strength of fully segmental beams with hybrid tendons was 30%

less than that of the monolithic beam with hybrid tendons. Due to a high concentration of rotation and deflection at

individual joints, the flexural strength of the partially prestressed segmental beam with hybrid tendons was 12.8% less

than that of the fully prestressed segmental beam with hybrid tendons.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

891

Despite this interest, no one to the best of our knowledge has studied the post-fire performance of internally post-

tensioned segmental concrete beams under monotonic static loading. Accordingly, this paper seeks to address the

behavior of such structural concrete beams in an attempt to investigate the effect of the length of the individual

segment, the number of joints between individual segments, and the heating temperature to which the member is

exposed to on the behavioral performance at serviceability and failure stages.

The structure of this article was designed in such a way to achieve the objectives of the research program that

starting with the experimental program, measurement of the important deformability aspects, comparisons and

interpretations analysis of experimental results, and highlighting the important conclusions.

2. Experimental Program

The experimental program consisted of twelve simply-supported post-tensioned segmental concrete beams. All

segmental beams were categorized into three groups depending on the number of the incorporated precast concrete

segments. All the PSC beams were designed and fabricated with a square cross-sectional configuration of 400 x 400

mm dimensions and 3150 mm overall length. In the first group, the PSC beams consisted of nine segments each of 350

mm length. While the second and third groups included seven and five segments each of 450 mm and 630 mm,

respectively.

Figure 1 shows the schematic configuration of the tested PSC beams while Figure 2 illustrates the details of the two

ends of the precast concrete segments in their opposite contact surfaces. Four concrete bulges, which play the role of

shear keys, and four cavities were fabricated at first and the second end of each segment, respectively. To achieve

perfect accommodation for the concrete shear key at the interface section, the dimensions of the cavity and the bulge

were adopted identical. The steel cages in each segment included longitudinal and transverse reinforcement of 8 mm

diameter deformed bars with yielding and ultimate strengths of 486 and 640 MPa, respectively (Figure 3). Twelve ∅ 8

mm longitudinal bars and ∅ 8 @ 60 mm c/c steel stirrups were used in each segment. The fabrication of the PSC

beams performed in two stages. In the first stage, all the precast concrete segments were cast in special metal forms

and continuously moist cured by wet burlap for seven days to achieve a concrete compressive strength of 40 MPa at

28-day for cube samples of 150 x 150 x 150 mm dimensions. Figure 4 shows the grading curves for the fine and coarse

aggregate that tested according to the ASTM C33-18 [17]. Consequently, in the second stage, the concrete segments

were assembled together according to the assigned groups using one eccentric prestressing force of 120 kN. To create

this force, one 12.7 mm diameter low-relaxation seven-wire steel strand (Grade 270, i.e., 1860 MPa) was used through

a plastic duct that had been fixed before casting with the steel cage. The distance from the center of the steel strand to

the soffit of the beam was considered to be equal to 120 mm in all test specimens. The jacking force was applied from

one end, where the adopted value was selected in such a way that to conform to the upper limit recommended by the

ACI 318M-14 Code. The test variables included three different structural systems:

• Post-tensioned segmental concrete beams with nine segments (PSC-9);

• Post-tensioned segmental concrete beams with seven segments (PSC-7);

• Post-tensioned segmental concrete beams with five segments (PSC-5).

For each case, three different degrees of heating temperature were selected (300, 500, or 700 °C). Table 1 lists the

test variables which have been considered. The abbreviation of each beam takes the format of the post-tensioned

segmental concrete beam - the number of precast concrete segments – the heating temperature. Except for the

reference beams PSC-9-REF, PSC-7-REF, and PSC-5-REF, the test program was carried out on each PSC beam

within two stages, mainly, the first stage during which the experimental beams were exposed directly to fire flame

and consequently followed by the second stage during which the beams were subjected to external monotonic static

loading up to failure. It is worth to mention that the reference beams were subjected to the monotonic static loading

only.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

892

Figure 1. Schematic configuration of the tested PSC beams (All dimensions are in mm)

Table 1. The adopted test variables of the experimental program

Heating temperature, °C Length of segment, mm Number of segments Specimen ID Group

-

350 9

PSC-9-REF

I 300 PSC-9-300

500 PSC-9-500

700 PSC-9-700

-

450 7

PSC-7-REF

II 300 PSC-7-300

500 PSC-7-500

700 PSC-7-700

-

630 5

PSC-5-REF

III 300 PSC-5-300

500 PSC-5-500

700 PSC-5-700

Civil Engineering Journal Vol. 6, No. 5, May, 2020

893

Figure 2. Details of precast concrete segments (All dimensions are in mm)

Figure 3. Reinforcement details of precast concrete segments (All dimensions are in mm)

Figure 4. Grading curves for fine and coarse aggregate tested according to the ASTM C33-18 [17]

0

20

40

60

80

100

120

0 10 20 30 40

Per

cen

tage

Pa

ssin

g (

%)

Sieve Size (mm)

Coarse Aggregate

Upper Limit

Lower Limit

0

20

40

60

80

100

120

0 2 4 6 8 10

Per

cen

tage

Pass

ing

(%

)

Sieve Size (mm)

Fine Aggregate

Upper Limit

Lower Limit

Civil Engineering Journal Vol. 6, No. 5, May, 2020

894

2.1. Stage I – Fire Test under Direct Flame Exposure

Three PSC beams in each group were exposed to a high temperature of (300, 500, or 700 °C) and uniformly

distributed loading of 3.22 kN/m, which simulated the dead load on the tested member, using 19 concrete blocks each

of 50 kg. This predetermined superimposed loading was applied before starting the fire test, which comprised a

heating and cooling phases. This load was maintained for the entire fire test. It is important to note that the fire

chamber was designed in such a way to allow expose the test beams to fire from three sides because a fire on three

sides is more common for beams (Figures 5 and 6). The heating temperature was controlled by changing the amount of

the supplied methane gas. The temperature readings from three thermocouples, Type K (Nickel-Chromium / Nickel -

Alumel) which can be used at temperatures up to 1100°C, was recorded during heating and cooling phases by a digital

thermometer reader. Thermocouples were installed in three locations at the center, the right and the left ends of the fire

chamber. The required time in minutes to attain the target temperature of 300, 500 and 700°C was 5, 7, and 12,

respectively. The heating phase of the fire test was continued for 60 minutes after the target temperature was achieved.

After that, the heating phase was terminated and the fire chamber allowed cooling through a cooling phase, which

performed gradually in the ambient air condition by removing the top cover of the chamber. It is worth to mention that

the heating phase followed the time-temperature curve provided in ASTM-E119 [18]. Accordingly, the test segmental

beams were exposed directly to fire flame in the presence of the prestressing force and the sustained superimposed

load. During the heating and cooling phases of the fire test, the mid-span section displacement due to the constant dead

load, the applied prestressing force, and heating temperature of the fire test was measured using a dial gauge of 0.002

mm/div. sensitivity.

Figure 5. Close-up view of the test under direct fire exposure

Civil Engineering Journal Vol. 6, No. 5, May, 2020

895

Figure 6. Setup of the fire test under direct flame exposure (All dimensions are in mm)

2.2. Stage II – Test under Monotonic Static Loading

To study the behavior and the residual ultimate strength for PSC beams after the cooling phase of the fire test, all

specimens including the reference beams at the ambient temperature were loaded to failure using a force control

module with a loading step of 2.5 kN in four-point bending using two symmetrical concentrated loads applied at

middle third of the span length (Figure 7).

-1 فصل

-2 فصل

-3 فصل

-4 فصل

-5 فصل

-6 فصل

-7 فصل

-8 فصل

-9 فصل

Figure 7. Close-up and schematic view of the test under the monotonic static loading

Civil Engineering Journal Vol. 6, No. 5, May, 2020

896

3. Experimental Results and Discussion

A summary of the test results including the test condition, the applied load at the collapse of the test beam, the

camber at the midspan section, the deflection of the midspan section due to the superimposed load, and the failure

mode are given for each beam. A detailed discussion for the test scenario is given in the following subsections.

3.1. Displacement of the Test Beams Prior to Fire Test

The combined effect of the applied prestressing force together with the beam self-weight caused upward deflection

(camber) at the midspan section before the application of the superimposed dead and the live load (Table 2). The PSC

beam’s midspan camber was measured by a mechanical dial gauge of 0.002 mm/division. Table 2 shows a significant

reduction in the midspan camber with the decreasing of the number of the incorporated segments of the member. It

should be noted that as the number of the incorporated segments decreased from nine (i.e., in Group I) to seven (i.e., in

Group II) or five (i.e., in Group III), the camber was decreased by 21% or 38%, respectively. The reason behind this

evidence is that the eccentric prestressing force which generates prestressing moment that causes more concentration

of rotations at the interface section between different segments due to the degradation of the flexural stiffness at these

sections. Accordingly, as more segments will be incorporated in the structural member as more concentration of

rotations and consequently displacements will occur. It should be noted that prior to exposing the test beams to the fire

test, a uniformly distributed load was applied, which simulated the superimposed load of the member itself. Each

specimen was load by 19 concrete blocks acting over the compression face of the member sections along its main axis.

This applied load sustained acting over the test beam during the first and second stages of the testing program until

failure. The measured midspan deflection due to the superimposed load is listed in Table 2. The same observation was

recorded for the values of midspan deflection which highly dependent on the number of the incorporated concrete

segments in the test beams.

Table 2. Deformability of the tested PSC beams due to prestressing force, self-weight, and superimposed load

Specimen

Camber due to prestressing force and self-

weight

Relative camber

value for group 𝜹𝒊, % Deflection due to

superimposed load, mm

Net camber before

fire test ∆𝒃𝒇, mm Camber for member

∆, mm

Average camber

for group ∆𝒊, mm 𝜹𝒊 = (

∆𝒊

∆𝑰

) × 𝟏𝟎𝟎

Gro

up

I

PSC-9-REF -2.9

∆𝐼 -2.9 100

- -2.90

PSC-9-300 -2.8 +0.36 -2.44

PSC-9-500 -3.1 +0.36 -2.74

PSC-9-700 -3.0 +0.36 -2.64

Gro

up

II

PSC-7-REF -2.3

∆𝐼𝐼 -2.3 79

- -2.30

PSC-7-300 -2.2 +0.27 -1.93

PSC-7-500 -2.4 +0.27 -2.13

PSC-7-700 -2.2 +0.27 -1.93

Gro

up

III

PSC-5-REF -1.8

∆𝐼𝐼𝐼 -1.8 62

- -1.80

PSC-5-300 -1.6 +0.20 -1.40

PSC-5-500 -1.8 +0.20 -1.60

PSC-5-700 -1.9 +0.20 -1.70

3.2. Displacement of the Test Beams During the Fire Test

During the heating and cooling phases of the fire test, the midspan section displacement was monitored

systematically for all beams. Results during these phases for all beams are summarized in Table 3. Camber versus time

history for the PSC beams in Groups I, II and III, respectively are shown in Figures 8 to 10. From these figures, it can

be seen that the increasing temperature during the heating phase led to a steep increase in the beam camber.

Meanwhile, a gradual decreasing of the achieved camber was observed during the cooling phase. In other words,

exposing a concrete beam to high temperature under sustaining eccentric load (i.e., prestressing force) led to a

progressive increase of the camber as a result of the degradation of the flexural stiffness of the member. The structural

response of the segmental beams during the heating phase of the fire test characterized by two-time intervals. In the

first interval (i.e., the progress of heating temperature interval), the increase of the camber values resulted essentially

from the generation of thermal strains caused by high thermal gradients. In the second time interval (i.e., the

stabilization of the heating temperature interval), as the temperature increased in the internal layers of concrete,

camber increased significantly and attained the peak value at the end of this time interval due to the reduction of the

thermal concrete strength due to the formation of the internal microcracks. The interference of the thermal effect on

different internal layers attributed to the degradation of the strength and the modulus of elasticity of concrete more

Civil Engineering Journal Vol. 6, No. 5, May, 2020

897

than that of the steel strand then camber increases at a high pace. In addition, camber increment is essentially attributed

to the high creep strains subsequent from high temperatures in concrete which compose of different materials and

strands. It is interesting from Figures 8, 9 and 10 to note that during the cooling phase of the fire test, the camber-time

diagram starts to descend. The length and the slope of the descending branch depend on the heating temperature that

the structural member was experienced and the number of incorporated concrete segments that the member was

composed of. The main reasons behind the appearance of the descending branch of the camber-time diagram are the

progressive increase of the prestressing losses due to the temperature difference (i.e., the temperature difference

between the prestressing steel in the heating zone and the device that receives the prestressing force (anchorages)

during concrete heating) and the creep of concrete that highly depend on the value of the heating temperature.

Table 3. Deformability of the PSC beams under fire tests

Specimen

Net

ca

mb

er b

efo

re f

ire t

est

∆𝒃

𝒇, m

m

Heating phase

Cooling phase

Net

ca

mb

er a

fter f

ire t

est

∆𝒂

𝒇, m

m

Rela

tiv

e r

esi

du

al

ca

mb

er

va

lue (

(∆𝒂

𝒇−

∆𝒃

𝒇)

∆𝒃

𝒇⁄

𝟏𝟎

𝟎 %

Ca

mb

er c

ha

ng

e r

ela

tiv

e

to r

efe

ren

ce m

em

ber, %

Ca

mb

er c

ha

ng

e r

ela

tiv

e

to n

ine-s

eg

men

t m

em

ber,

%

Progress of

heating temp.

Stabilization of

heating temp.

Tim

e p

erio

d,

min

Ca

mb

er

va

ria

tio

n, m

m

Tim

e p

erio

d,

min

Ca

mb

er

va

ria

tio

n, m

m

Tim

e p

erio

d,

min

Dis

pla

cem

en

t

va

ria

tio

n, m

m

PSC-9-REF -2.90 - - - - - - -2.90 - - -

PSC-9-300 -2.44 0-5 -0.2 5-65 -2.0 65-245 +0.9 -3.74 +53 +29 -

PSC-9-500 -2.74 0-7 -0.5 7-67 -3.3 67-380 +1.9 -4.64 +69 +60 -

PSC-9-700 -2.64 0-12 -1.0 12-72 -6.7 72-520 +4.1 -6.24 +136 +115 -

PSC-7-REF -2.30 - - - - - - -2.30 - - -

PSC-7-300 -1.93 0-5 -0.1 5-65 -1.6 65-245 +0.7 -2.93 +52 +27 -22

PSC-7-500 -2.13 0-7 -0.3 7-67 -2.9 67-380 +1.6 -3.73 +75 +62 -20

PSC-7-700 -1.93 0-12 -0.9 12-72 -6.2 72-520 +3.9 -5.13 +166 +123 -18

PSC-5-REF -1.80 - - - - - - -1.80 - - -

PSC-5-300 -1.4 0-5 -0.07 5-65 -1.27 65-245 +0.67 -2.07 +48 +15 -45

PSC-5-500 -1.6 0-7 -0.2 7-67 -2.3 67-380 +1.2 -2.90 +81 +61 -38

PSC-5-700 -1.7 0-12 -0.8 12-72 -5.2 72-520 +3.4 -4.30 +153 +139 -31

In Table 3, the accumulative value for the midspan section displacement including the net camber before the fire

test ∆𝑏𝑓, the camber variation during the heating phase and the displacement variation during the cooling phase is

called the net camber after the fire test ∆𝑎𝑓. The difference between ∆𝑎𝑓 and ∆𝑏𝑓 indicates the size of the residual

deformation that the test member was experienced due to the temperature exposure. Depending on the number of the

incorporated concrete segments in the test beams and the heating temperature of the fire test, the relative residual

midspan camber consisted (48 – 53%), (69 – 81%), and (136 – 166%) for heating temperature of 300, 500, and 700

°C, respectively. Obviously after the heating and cooling phases of the fire test, with the increasing of the heating

temperature, the net value for camber was increased in all test beams. At the end of the fire test, the net camber values

for beams of Group I that exposed to different heating temperatures were increased by 29, 60, and 115% compared to

the net camber value of the reference beam PSC-9-REF for heating temperatures of 300, 500, and 700 °C,

respectively. Meanwhile, this increase in the net camber value was attained 27, 62, and 123% for specimens of Group

II and 15, 21, and 139% for specimens of Group III for the same mentioned above temperatures compared to their net

camber values for reference beams, respectively, PSC-7-REF and PSC-5-REF. From Table 3 and Figures 8 to 10, it

should be noticed that the deformability of the test member depends on the heating temperature that the member was

exposed to and the length of the time interval of the fire test. The camber was increased during the heating period and

reached its maximum value at the end of the stabilization period of the heating temperature. Table 3 and Figures 11 to

13 illustrate the comparison of the camber – time history during the fire test for different test beams depending on the

number of the incorporated concrete segments. Obviously at the same heating temperature of the fire test, as the

number of the incorporated concrete segments decreased the net camber value at the end of the fire test was decreased.

In comparison to the test beams with nine concrete segments, it can be noted that at a heating temperature of 300 °C

the net camber value decreased by 22 and 45% for beams with seven and five concrete segments, respectively.

However, at 500 °C, the reduction of the net camber value attained 20 and 38% for test beams with seven and five

concrete segments, respectively. The same behavior was monitored at the fire test of 700 °C heating temperature,

whereas the reduction of the net camber value reached 18 to 31% for the same mentioned above specimens,

respectively. This behavior of deformability can be interpreted by the fact that, as the number of dry joints increased

Civil Engineering Journal Vol. 6, No. 5, May, 2020

898

0

2

4

6

8

10

12

0 100 200 300 400 500 600

Cam

ber,

mm

Time, min

PSC-7-300

PSC-7-500

PSC-7-700

0

2

4

6

8

10

12

0 100 200 300 400 500 600

Cam

ber ,

mm

Time, min

PSC-9-300

PSC-7-300

PSC-5-300

the rotational displacement between different concrete segments due to the creep of concrete under the dual effect of

prestressing force and fire exposure also increased which in turn achieved higher camber values accordingly.

Additionally, the comparisons illustrated in Figures 8 to 10, also in Figures 11 to 13 show that for the same number of

segments, as the heating temperature increased the camber of the tested member was increased. Meanwhile for the

same heating temperature, as the number of concrete segments increased the camber of the tested specimen was also

increased.

Figure 8. Camber - time history for PSC beams of Group I at different heating temperatures

Figure 9. Camber - time history for PSC beams of Group II at different heating temperatures

Figure 10. Camber - time history for PSC beams of Group III at different heating temperatures

Figure 11. Camber-time history of PSC beams depending on the number of incorporated segments at fire test of 300˚C

0

2

4

6

8

10

12

0 100 200 300 400 500 600

Cam

ber ,

mm

Time, min

PSC-9-300

PSC-9-500

PSC-9-700

0

2

4

6

8

10

12

0 100 200 300 400 500 600

Cam

ber ,

mm

Time, min

PSC-5-300

PSC-5-500

PSC-5-700

Civil Engineering Journal Vol. 6, No. 5, May, 2020

899

0

2

4

6

8

10

12

0 100 200 300 400 500 600

Cam

ber ,

mm

Time, min

PSC-9-500

PSC-7-500

PSC-5-500

0

2

4

6

8

10

12

0 100 200 300 400 500 600

Cam

ber ,

mm

Time, min

PSC-9-700

PSC-7-700

PSC-5-700

Figure 12. Camber-time history of PSC beams depending on the number of incorporated segments at fire test of 500˚C

Figure 13. Camber-time history of PSC beams depending on the number of incorporated segments at fire test of 700˚C

3.3. Displacement of the Test Beams under Monotonic Static Loading

To investigate the structural behavior of the segmental beams after fire exposure and to evaluate their post-fire

resistance, all test beams were subjected to monotonic static loading under the effect of two concentrated loads that

were applied at a distance of 975 mm from the nearest support for each.

The load was applied gradually up to failure, where each load increment consisted of 2.5 kN. It is worth to mention

that after fire testing microcracks with different intensity were observed on the surfaces of the test specimens. The

intensity and distribution of these cracks depended on the heating temperature that the structural member was exposed.

The load-deflection curves for PSC beams of Groups I, II and III, respectively, are shown in Figures 14 to 16. Based

on these figures it can be observed that, for specimens with the same number of incorporated concrete segments, as the

heating temperature increased the load capacity of the test specimen was significantly decreased. This fact can be

attributed to the considerable microcracking that formed during the exposure to elevated temperatures. Accordingly, as

the heating temperature increases the microcracking process progressively increases. The formed microcracks

affected the bond between the aggregate and the cement paste that caused a reduction of the effective section area and

the concrete modulus of elasticity which in turn led to significant degradation of the overall stiffness of the test

member. Figures 14 to 16 show also that, the effect of the heating temperature of 300 °C was very limited on changing

the load-deflection behavior of segmental beams and on the reduction of their load-carrying capacity. Meanwhile, as

the heating temperature during the fire test was increased to 500 or 700 °C, the load-deflection behavior became

flattered and the reduction of the load capacity became more than the reduction at 300 °C due to the same reasons that

mentioned above. It is worth mentioning that in segmental concrete beams the lower concrete chord does not

contribute to the flexural resistance of the member due to the dry joints created between different concrete segments.

As a result, the externally applied load in this system of construction (i.e., post-tensioned segmental concrete beams)

resisted by the internal couple represented by the lower tension in the prestressing steel and the upper compression

forces resultant in the top concrete chord. From the flexural resistance point of view, the concrete in the lower chord is

playing just the role of a protector layer from fire and environmental attacks. The applied load and the corresponding

midspan deflection for all PSC beams at different stages of loading are listed in Table 4. In this table, three different

loads were adopted to compare the performance of all test beams, mainly, the applied load of 27.3 kN which

represents 65% of the minimal failure load among all test beams, the service load of the nine-segment members which

considered equal to 65% of the failure load of the corresponding nine-segmental beam, and the failure load of the

corresponding beam. As shown in this table and illustrated in Figures 14 to 16, the deflection due to the applied load

was increased with the increasing of the heating temperature that the structural member was exposed during the fire

test. At the applied load of 27.3 kN, the deflection increase relative to the reference member (i.e., member which was

Civil Engineering Journal Vol. 6, No. 5, May, 2020

900

0

25

50

75

100

125

0 5 10 15 20 25 30 35 40

Lo

ad

, k

N

Deflection, mm

PSC-9-REF PSC-9-300 PSC-9-500 PSC-9-700

0

25

50

75

100

125

0 5 10 15 20 25 30 35 40

Lo

ad

, k

N

Deflection, mm

PSC-7-REF PSC-7-300 PSC-7-500 PSC-7-700

0

25

50

75

100

125

0 5 10 15 20 25 30 35 40

Lo

ad

, k

N

Deflection, mm

PSC-5-REF

PSC-5-300

PSC-5-500

PSC-5-700

not exposed to fire test), for PSC beams of different numbers of incorporated concrete segments 9, 7, and 5 that

exposed to heating temperature of 300 ºC, was 44, 31, and 20%, respectively. This increase of the midspan deflection

for specimens that exposed to 500 ºC attained 134, 105, and 80%, respectively. It is important to note that the worst

case was for the segmental beams that exposed during the fire test to the heating temperature of 700 ºC. For these

beams, the overall stiffness was highly affected by this range of temperature and the relatively long time period of

exposure. The midspan deflection increase reached 480, 405, and 310 % for the specimens of 9, 7, and 5 concrete

segments, respectively, in comparison to their reference beams.

Figure 14. Load-deflection curve of Group I

Figure 15. Load-deflection curve of Group II

Figure 16. Load-deflection curve of Group III

The effect of the number of the incorporated concrete segments on the behavior of the test specimens was

demonstrated through the load-midspan deflection diagrams during the monotonic static loading test, see Table 4 and

Figures 17 to 20. Obviously, the overall stiffness of the reference beams and the beams that exposed to the fire test of

different heating temperatures was highly depending on the number of the incorporated concrete segments.

Accordingly, as this number increased the overall stiffness decreased and in turn the midspan deflection increased.

Needless to say that the reason behind this fact is as the number of the incorporated in the structural member concrete

segments increases, the possibility of the rotational displacement in the created interface sections (i.e., joints) between

Civil Engineering Journal Vol. 6, No. 5, May, 2020

901

0

25

50

75

100

125

0 5 10 15 20 25 30 35 40

Lo

ad

, k

N

Deflection, mm

PSC-5-REF

PSC-7-REF

PSC-9-REF

0

25

50

75

100

125

0 5 10 15 20 25 30 35 40

Lo

ad

, k

N

Deflection, mm

PSC-5-300

PSC-7-300

PSC-9-300

different segments also increases which leads to excessive translational displacement. Accordingly, at a load equals to

the service load of nine-segment beams the midspan deflections for seven-segment and five-segment beams were

averagely decreased by about 37 and 63%, respectively, at different heating temperatures, see Table 4. It is very

interesting to note that the difference in performance, under the applied monotonic static loading depending on the

number of the incorporated concrete segments, was decreased as the heating temperature of the fire test was increased.

The reason behind this observation was the huge deterioration and microcracking intensity that occurred in concrete

composition at high temperatures (i.e., 500 and 700 ºC) regardless of the number of the incorporated segments. This

fact led to minimizing the difference between the load-midspan deflection behavior for the members that exposed to

the same heating temperature (i.e., 500 or 700 ºC) but consisted of different numbers of concrete segments (i.e., 9, 7 or

5).

Table 4. Deflections at different loading stages of test beams

Specimen ID

At a load of 27.3 kN At service load of nine-segment members which

adopted 65% of failure load At failure load

Deflection,

mm

Deflection increase

relative to reference

member, %

Service load of

nine-segment

specimen, kN

Deflection,

mm

Deflection change

relatie to nine-

segment member, %

Failure

load, kN

Deflection,

mm

Deflection increase

relative to reference

member, %

PSC-9-REF 0.86 - - - - 86 29 -

PSC-9-300 1.24 44 53.3 4.57 - 82 33 14

PSC-9-500 2.01 134 45.5 4.71 - 70 34 17

PSC-9-700 4.99 480 27.3 4.99 - 42 39 34

PSC-7-REF 0.62 - - - - 98 30 -

PSC-7-300 0.81 31 53.3 2.84 -38 92 31 3

PSC-7-500 1.27 105 45.5 2.95 -37 77 33 10

PSC-7-700 3.13 405 27.3 3.13 -37 47 35 17

PSC-5-REF 0.49 - - - - 108 20 -

PSC-5-300 0.59 20 53.3 1.61 -65 101 20 0

PSC-5-500 0.88 80 45.5 1.77 -63 83 22 10

PSC-5-700 2.01 310 27.3 2.01 -60 51 25 25

Figure 17. Load-midspan deflection curves for reference PSC beams depending on the number of the incorporated concrete

segments

Figure 18. Load-midspan deflection curves for PSC beams, which exposed during fire test to 300 ˚C heating temperature,

depending on the number of the incorporated concrete segments

Civil Engineering Journal Vol. 6, No. 5, May, 2020

902

0

25

50

75

100

125

0 5 10 15 20 25 30 35 40

Lo

ad

, k

N

Deflection, mm

PSC-5-500

PSC-7-500

PSC-9-500

0

25

50

75

100

125

0 5 10 15 20 25 30 35 40

Lo

ad

, k

N

Deflection, mm

PSC-5-700

PSC-7-700

PSC-9-700

Figure 19. Load-midspan deflection curves for PSC beams, which exposed during fire test to 500 ˚C heating temperature,

depending on the number of the incorporated concrete segments

Figure 20. Load-midspan deflection curves for PSC beams, which exposed during fire test to 500 ˚C heating temperature,

depending on the number of the incorporated concrete segments

3.4. Load-carrying Capacity and Failure Modes of the Test Beams

Table 5 shows the outcome data for the failure load of all test specimens. Obviously, the exposure of the post-

tensioned segmental concrete beams to high heating temperatures during the fire test stage affected seriously the post-

fire performance and the load-carrying capacity of the mentioned structural members during their exposure to a

monotonic static loading stage. Test results revealed that two parameters were significantly influenced the load

capacity of the investigated structural members, mainly, the heating temperature value that the member was exposed

during the fire test stage and the number of the incorporated concrete segments that the member was consisted of.

Increasing each or both of these parameters resulted in a reduction of the load-carrying capacity due to the reasons

discussed above. In Group I of nine-segment beams, the residual load-carrying capacity was 95, 81 and 49% of the

failure load of the reference beam PSC-9-REF for members exposed during fire test to heating temperatures of 300,

500, and 700 ºC, respectively. The reduction of the failure load for this group in comparison to the reference beam was

ranged between 5 to 51%. Whereas in Group II of seven-segment beams and in Group III of five-segment beams, the

failure load was 94, 79, and 48 % of the load capacity of the reference beam PSC-7-REF and 94, 77, and 47% of the

failure load of the reference beam PSC-5-REF for specimens that experienced during fire test stage to temperatures of

300, 500, and 700 ºC, respectively. Accordingly, the reduction of the load capacity for Groups II and III in comparison

to their reference beams was ranged between 6 to 52% and 6 to 53%, respectively.

For the three groups when the heating temperature increased up to 300 ºC, the average residual strength decreased

by a rate of 6% compared to the reference beams. Increasing during the fire test stage the heating temperature up to

500 ºC led to a decrease in the average residual strength by 21%. Meanwhile, the average residual strength decreased

by a rate of 52% when the temperature reached 700 ºC compared to the reference beams. In comparison to the test

beams with nine concrete segments, it can be noted that at a heating temperature of 300 °C the failure load value

increased by 12 and 23% for beams with seven and five concrete segments, respectively. However, at 500 °C, the

increase of the failure load value attained 10 and 19% for test beams with seven and five concrete segments,

respectively. The same behavior was monitored for the members that exposed to a fire test of 700 °C heating

temperature, whereas the increase of the load-carrying capacity reached 12 to 21% for the same mentioned above

specimens, respectively. Figure 21 shows the comparison of the load-carrying capacity between different test beams.

Figure 22 demonstrates the cracking distribution and the failure mode of the test specimens. It should be mentioned

that in all test specimens the mode of failure included the crushing of the concrete in the compression zone of the

central concrete segment due to the excessive rotation that occurred during the application of the external loading.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

903

0

20

40

60

80

100

120

0 100 200 300 400 500 600 700 800

Lo

ad

-Carryin

g C

ap

acit

y,

kN

Heating Temperature, ºC

Group I

Group II

Group III

Also, the joint-opening width at failure stage was highly depended on the heating temperature that the post-tensioned

segmental concrete member was exposed to.

Table 5. The load-carrying capacity of test PSC beams

Specimen Failure Load,

(kN)

Change of failure load relative

to a reference member, %

Change of failure load relative to

nine-segment member, %

PSC-9-REF 86 - -

PSC-9-300 82 -5 -

PSC-9-500 70 -19 -

PSC-9-700 42 -51 -

PSC-7-REF 98 - +14

PSC-7-300 92 -6 +12

PSC-7-500 77 -21 +10

PSC-7-700 47 -52 +12

PSC-5-REF 108 - +26

PSC-5-300 101 -6 +23

PSC-5-500 83 -23 +19

PSC-5-700 51 -53 +21

Figure 21. Comparison of the load-carrying capacity between different groups

PSC-5-REF PSC-7-REF

Civil Engineering Journal Vol. 6, No. 5, May, 2020

904

PSC-9-REF PSC-5-300

PSC-7-300 PSC-9-300

PSC-5-500 PSC-7-500

PSC-9-500 PSC-5-700

PSC-7-700 PSC-9-700

Figure 22. The failure mode of test beams

Civil Engineering Journal Vol. 6, No. 5, May, 2020

905

4. Conclusion

It was observed that decreasing the number of the incorporated segments of the beam led to a significant reduction

in the midspan camber (i.e., decreasing the number of segments from nine to seven or five, which led to a decrease in

the camber by 21% or 38%, respectively). The structural response of the segmental beams during the heating phase

was characterized by two-time intervals. In the first interval, the increase of the camber resulted from the thermal

strains caused by high thermal gradients. While the camber was progressively increased and attained its peak value at

the end of the second interval as the temperature increased in the internal layers of concrete due to the migration of the

heat energy across the member that deteriorated the texture of the concrete and caused microcracking of larger surface

areas. Under the cooling event, the camber-time diagram descended where the length and the slope of the descending

branch depend on the heating temperature and the number of incorporated concrete segments. The main reasons

behind the appearance of the descending branch are the progressive increase of the prestressing losses due to the

temperature difference and the creep of concrete that highly depends on the value of the heating temperature.

The relative residual midspan camber at the end of the cooling phase consisted (48–53%), (69–81%), and (136–

166%) for heating temperature of 300, 500, and 700 °C, respectively, depending on the number of the incorporated

concrete segments in the test beams and the heating temperature of the fire test. For beams with the same number of

incorporated concrete segments, as the heating temperature increased during the fire event, the load-carrying capacity

of the test specimen was significantly decreased. The reduction of the failure load for the nine-segment beams in

comparison to the reference beam was ranged between (5-51%), whereas in the seven-segment beams and the five-

segment beams, the reduction attained (5 –52%) and (6–53%), respectively. As the heating temperature increased up

to 300 ºC, the average residual strength decreased by a rate of 6% compared to the reference beams. Increasing during

the fire test stages the heating temperature up to 500 ºC led to a decrease in the average residual strength by 21%.

Meanwhile, the average residual strength decreased by a rate of 52% when the temperature reached 700 ºC compared

to the reference beams. The mode of failure included the crushing of the concrete in the compression zone of the

central concrete segment.

5. Conflicts of Interest

The authors declare no conflict of interest.

6. References

[1] Al-Gorafi, M A, A A A Ali, I Othman, M S Jaafar, and M P Anwar. “Externally Prestressed Monolithic and Segmental

Concrete Beams under Torsion: a Comparative Finite Element Study.” IOP Conference Series: Materials Science and

Engineering 17 (February 1, 2011): 012041. doi:10.1088/1757-899x/17/1/012041.

[2] Chan, Y.N., X. Luo, and W. Sun. “Compressive Strength and Pore Structure of High-Performance Concrete after Exposure to

High Temperature up to 800°C.” Cement and Concrete Research 30, no. 2 (February 2000): 247–251. doi:10.1016/s0008-

8846(99)00240-9.

[3] Koksal, F, O Gencel, W Brostow, and H E Hagg Lobland. “Effect of High Temperature on Mechanical and Physical Properties

of Lightweight Cement Based Refractory Including Expanded Vermiculite.” Materials Research Innovations 16, no. 1

(February 2012): 7–13. doi:10.1179/1433075x11y.0000000020.

[4] Aslani, Farhad. “Prestressed Concrete Thermal Behaviour.” Magazine of Concrete Research 65, no. 3 (February 2013): 158–

171. doi:10.1680/macr.12.00037.

[5] Myers, John J., and Wendy L. Bailey. "Seven-Wire Low Relaxation Prestressing Tendon Subjected to Extreme Temperatures:

Residual Properties." International Journal of Engineering Research and Science and Technology 4, no. 3 (March 2015): 223-

239.

[6] Abdelrahman, A., Nofel, N., Ghalib, A., El-Afandy, T. “Behaviour of Prestressed Concrete Beams Subjected to Fire.” Housing

and Building National Research Centre Journal 7, no. 2 (August 2011): 38-55.

[7] Bennetts, I., and W. South. “Real Fire Test on Concrete Columns and Post-Tensioned Slabs.” Fire Safety Science 11 (2014):

558–571. doi:10.3801/iafss.fss.11-558.

[8] Izzet, Amer Farouk. "Effect of High Temperature (Fire Flame) on the Behavior of Post-tensioned Concrete Beams."

Association of Arab Universities Journal of Engineering Sciences 25, no. 3 (2018): 49-68.

[9] Phan, Long T., and Nicholas J. Carino. “Fire Performance of High Strength Concrete: Research Needs.” Advanced Technology

in Structural Engineering (April 27, 2000). doi:10.1061/40492(2000)181.

[10] Zhang, Li, Ya Wei, Francis T. K. Au, and Jing Li. “Mechanical Properties of Prestressing Steel in and after Fire.” Magazine of

Concrete Research 69, no. 8 (April 2017): 379–388. doi:10.1680/jmacr.15.00267.

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[11] Sivaleepunth, Chunyakom, Junichiro Niwa, Dinh Hung Nguyen, Tsuyoshi Hasegawa, and Yuzuru Hamada. “Shear Carrying

Capacity of Segmental Prestressed Concrete Beams.” Doboku Gakkai Ronbunshuu E 65, no. 1 (2009): 63–75.

doi:10.2208/jsceje.65.63.

[12] Nguyen, Dinh Hung, Ken Watanabe, Junichiro Niwa, and Tsuyoshi Hasegawa. “Modified Model for Shear Carrying Capacity

of Segmental Concrete Beams with External Tendons.” Doboku Gakkai Ronbunshuu E 66, no. 1 (2010): 53–67.

doi:10.2208/jsceje.66.53.

[13] Algorafi, M.A., A.A.A. Ali, I. Othman, M.S. Jaafar, and M.P. Anwar. “Experimental Study of Externally Prestressed

Segmental Beam under Torsion.” Engineering Structures 32, no. 11 (November 2010): 3528–3538.

doi:10.1016/j.engstruct.2010.07.021.

[14] Yuan, Aimin, Hangs Dai, Dasong Sun, and Junjun Cai. “Behaviors of Segmental Concrete Box Beams with Internal Tendons

and External Tendons Under Bending.” Engineering Structures 48 (March 2013): 623–634.

doi:10.1016/j.engstruct.2012.09.005.

[15] Moubarak, A., Kassem, N., Emad, E., Taher, S. “Eccentricity Shift in Externally Prestressed Segmental Concrete Beams.”

International Conference on Advances in Structural and Geotechnical Engineering, ICASGE’15, 6-9 (April 2015): Hurghada,

Egypt. pp 1-21.

[16] Jiang, Haibo, Qi Cao, Airong Liu, Tianlong Wang, and Yun Qiu. “Flexural Behavior of Precast Concrete Segmental Beams

with Hybrid Tendons and Dry Joints.” Construction and Building Materials 110 (May 2016): 1–7.

doi:10.1016/j.conbuildmat.2016.02.003.

[17] ASTM Standard C33/C33M. "Specification for Concrete Aggregates," ASTM International, West Conshohocken, PA, (2018).

[18] ASTM Standard E119-10. "Test Method for Fire Tests of Building Construction and Materials," ASTM International, West

Conshohocken, PA, (2010).

Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

907

Optimalization of the Ferronickel Production Process through

Improving Desulfurization Effectiveness

Izet Ibrahimi a, Nurten Deva

a*, Sabri Mehmeti

a

a Faculty of Geosciences; University of Mitrovica “Isa Boletini” Ukshin Kovaçica, 40000 Mitrovicë, Kosovo .

Received 19 December 2019; Accepted 22 March 2020

Abstract

Desulphurization of Ferronickel in the converters with oxygen is the most complex part of the technological process in

the Drenas foundry. Sulphur in the ferronickel melting is mostly in the form of FeS, with a melting temperature of

1195oC, and it has tendency to dissolve indefinitely in liquid iron. Our objective is to determine the sulphur removal

coefficient, as a key indicator of the desulphurization efficiency in the converter, by measuring the activity and

concentration of sulphur and other elements in liquid Fe and melting. Determination of this coefficient is done according

to the analytical method, while comparing the current process parameters with those of the new desulfurization methods,

other indicators of the refining process are determined. The refining process and the effective conduct of the study

depend on the XRD analysis database of metal and slag, and as well of the technological refining process analysis data.

Research has shown that desulfurization efficiency is a function of the sulphur removal coefficient, respectively; metal

composition, slag, oxygen activity, CaO/SiO2 ratio, sulphide capacity, fluidity, surface pressure, etc.). In addition to this

coefficient, other indicators of refining process optimization are defined.

Keywords: Ferronickel; Slag; Sulphur Portion Coefficient; Desulphurization; Sulphide Capacity; Refined Ferronickel.

1. Introduction

The pyrometallurgical obtaining of ferronickel from oxide-laterite ores, regardless of the degree of technical-

technological excellence even in the newer processes, has many unresolved technical and technological problems, first

of all the process of refining ferronickel in the converter is followed by low efficiency due to the lack of

desulphurization outside the furnace as well as the low desulphurization dynamics in the converter. During

pyrometallurgical processing the converter is one of the most important aggregates to produce and to refine

ferronickel. The oxygen blowing process is necessary to decrease the sulphur, phosphorus, carbon, silicon and the iron

content in the FeNi metal to the requested levels [1].

During the production of ferronickel in the electric furnace, in addition to iron and nickel, and other metals such as

cobalt, manganese, chromium, sulphur, copper, silicon, phosphorus, carbon, etc., pass to the alloy, which adversely

affects the properties of it and their removing presents additional difficulties, adversely affecting the process economy.

Since ferronickel, is used for the production of various steels must contain a minimum amount of sulphur below

0.04% [2, 3], then it is necessary to make the deepest desulphurisation, whether through out-furnace desulphurisation,

before that the metal passes to the converter or during process of refining in the converter. With technological

advancements, the requirement for minimum sulphur content in steels has reached up to 0.02%, while for special

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091516

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

908

steels up to 0.005%. Under the current operating conditions at the Foundry of NewCo Ferronickel in Drenas, such

quality parameters of refined ferronickel are almost inaccessible. Refining, and especially sulphur removal phase as it

is not being applied between the desulphurization phase out-furnace, but all the sulphur removal takes place in

converter, followed by a slow process and with a high consumption of materials and energy resources [4, 5].

Sulphur in the components of charge is pyrite, metallic and organic origin. This third form is likely to have the

greatest negative impact, since it is thermodynamically more stable and is expected to burn to the melting zone [6]. In

the electric furnace ferronickel, sulphur is most commonly present in the form of FeS, as iron constitutes the largest

proportion of molten ferronickel. During oxygen blowing in the converter, one part of the sulphur is removed with

slag and the other part with gas. Removal of ferrous sulphur is possible with the help of oxidizing slag [4]. The quality

of the refined ferronickel, the effectiveness of desulphurization and the stability of the technological parameters are

directly related to the sulphur separation coefficient, the physico-chemical properties of the metal as well as the

fluidity, optimal basicity and sulphide capacity of the slag [6, 7].

The main purpose of this study is to provide solutions that would improve the effectiveness of desulphurization, by

reducing the duration time of the metal in the converter, which would directly affect the optimization of refining

process in the converter and generally to decrease the production costs of ferronickel at the "Ferronickel Foundry" in

Drenas [8].

2. Research Methodology

This study is based on chemical analyzes of raw materials, electric furnace of ferronickel, chemical composition of

metal and slag according refining stages, gases, smelter and other technological materials. Based on these data and

process operational indicators, we have determined the sulphur separation coefficient as a key determinant of

desulphurization effectiveness. Also are determined the other performance indicators of the production process at the

"Ferronickel Foundry" in Drenas, according to the comparative method. Comparison of outputs, process parameters,

and production practices, are based on review the production period November 1997 and January 2009.Whereas the

effectiveness of desulphurization according to these operating parameters is compared with theoretical parameters,

assuming the development of off-furnace desulfurization, where magnesium would be used as a desulphurizer Stirling

[9].

Evaluation of desulphurization efficiency and measurement of process indicators are supported by chemical-

technological analysis of furnace metal and slag, metal and slag of converters, chemical and granulometric

composition of CaCO3, as and content and amount other technological materials, which are used during the

production process in the "Ferronickel Foundry" [10].

This study program will follow the standard research diagram are shown in Figure 1. Important indicators of the

experimental part are given in the load ratio see table 1. The charge of the refining process on the oxygen converter at

the Ferronickel Foundry in Drenas is designed depending on the chemical composition of the metal. An important

indicator in the determination of process parameters is the composition of Si and C in the furnace metal (EF). For our

case study, we converted 1.5 t of unrefined scrap and 1.5 t of refined scrap into the converter.

The scrap is used for thermal balancing and for improving the chemical environment in the process reaction area.

During the first phase (the burning process of Si and C) due to the high Si coefficient on the metal, the phase lasted 16

min, and the whole process would take place in an acidic environment with high heat release. For the thermal and

chemical balancing of the process, the amount and frequency of CaCO3 dosing will be increased. After this phase the

chemical analysis of the metal is carried out, on the basis of which the second stage of the process is designed -

refining under the basic environment. At this stage it is projected: the addition of CaO3 (kg / min), the flow of O2, (m3 /

h), the position of the spear towards the metal bath and the best possible removal of the process slag. During the third

phase besides the one mentioned above, other added additions in charge are calculated. Since at this stage high effects

of refractory material damage on the converter appear, the amount of MgO (90%) refractory brick powder mixed with

CaCO3 in 1: 2 ratios and the amount of refined metal scrap will be added, which would help enrich the metal with Ni

in FeNi. For successful conduct of research program besides samples for chemical analysis of metal we have also

analyzed 3 samples for chemical analysis of slag. The refined metal has been subjected to the deoxidation process. For

the development of this intermediate process, Fe-Si and Al have been used as a means of deoxidation.

Furthermore, the study relates to the experimental research, which used the results of XRD analysis by the NewCo

“Ferronickel” laboratory in Drenas, the values of the operating indicators in the converters unit, but also descriptive

data and comparative models from good practices of refining processes.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

909

Figure 1. Flowchart for the research methodology

Table 1. New Co Ferronikeli complex L.L.C Convertors – Charging Report

Serial No. of charge; (C=0844) No. of convertor.1 Lance: 1 Ladle for metal (Temp. of refractory = 700-800 0C)

O2 H2O

Flow (m3/h) Flow (m

3/h) -51.2

Convertor charge; 18 Oxygen indication (m3) No. of ladle.;18/4

P (bar) – 12.

(Flow O2 (mn3/min);

1. Phase 1 and 2 = 30 mn3/min

2. Phase 3,4,5 = 25 mn3/min

3. Phase 6,7 = 30 mn3/min

4. Phase 8, 9 = 35 mn3/min)

P (bar) – 10.1 T(H2O) = 16

0C

EF metal temperature

(TN); 1490 – 1500 0C

Before blowing - 874071

After blowing- 876021

Consumption- 1950

Departure time; 105

Tapping time of E.F.; 110

Arrival time; -135

Preheating time: 800 0C

Gross weight; 25 t

Dead weight; 2 t

Net weight; 23 t

No. of charge; 18

Phase Time

(from/to)

Duration

(min)

O2 – consumption

(mn3)

Lance

position

Scrap before charge

(t)Ni(t) CaCO3

(MgO + CaCO3

= 2:1) (kg)

Temp. by

phase (0C)

1 141

151

10 302 1.3 1.5 (t) from EF 460 120

2 154

202

6 (Analysis) 204 1.2

Scrap during charge 1.5 (t)

from conv.

600 100 1629

3 206

212

6 208 1.1 880 100

4 220

226

6 (Analysis) 218 1.2 950 60 1565

5 235

241

6 204 1.1 700 120

6 245

251

6 198 1.1 600 100 1500

7 258

304

6 213 1.1 850

8 315

320

5 201 1.1 700 1572

9 325

327

2 70 1.1 - 1568

10 330

333

3 110 - 1630

Total = 55 min 1942 3 (t) 5740

Civil Engineering Journal Vol. 6, No. 5, May, 2020

910

Table 2. Chemical analysis of Fe-Ni

Samples As Fe Si Cr C P S Ni Co Cu Temp. (0C)

Sample from EF 0.03 72.52 2,68 0.76 0.60 0.07 1,62 11.22 0.74 0.03 1629

Sample 1 After the seventh phase 0.67 15.63 1565

Sample 2

Sample 3

Final analysis 0.44 17.61 1568

Table 3. Chemical analysis of slag

CaO SiO2 FeO Cr2O3 Ni MgO Al2O3

Sample 1 23.26 18.40 36.18 2.86 0.19 10.10 0.93

Sample 2 23.37 17.42 34.81 2.94 0.21 9.79 0.93

Sample 3 19.91 15.14 45.95 1.00 0.25 11.60 0.78

In order to compare the heat balance with the chemical composition of the metal from the electric furnace and to

determine the duration of the process phases, have been measured in advance; the temperature and residence time of

the metal in ladle from the start of the electric furnace up to the spill to the converter, the amount of metal spilled into

the converter and scrap added before the start of the first phase.

After chemical analysis of charge were evaluated; elongation period of silicon oxidation, decarburization and

removal efficiency of other impurities from metal, consumables of refractory material, oxygen flow (Nm3 / min), spear

height (m) above metal melting, temperature (OC) afterwards at which stage of the process, the chemical composition

of the metal and coating after the first, second and third phases, the composition and amount of scrap and other

additives. After chemical analysis of the charge have been determined the ratios; CaO/SiO2, time required for

oxidation of silicon, carbon and other impurities, oxygen flow (Nm3/min), the spear position (m) above and

temperature (OC) for every each phase of process.

Ferronickel of electric furnace of "Ferronickel Foundry" in Drenas, is followed by various impurities, such as:

silicon, chromium, copper, carbon, sulfur, phosphorus, phosphorus, cobalt, etc. Sulfur, of all the impurities, has the

greatest negative impact on ferronickel properties, when is used for the production of steel, and its removal proceed

slowly and is followed by high energy costs.

3. Desulphurization of Ferronickel

Ferronickel of electric furnace of "Ferronickel Foundry" in Drenas, is followed by various impurities, such as:

silicon, chromium, copper, carbon, sulfur, phosphorus, phosphorus, cobalt, etc. Sulphur, of all the impurities, has the

greatest negative impact on ferronickel properties, when is used for the production of steel, and its removal proceed

slowly and is followed by high energy costs. Thus, this phase of the refining process is quite complex, which is usually

preceded by desulfurization in the ladie for metal or in the case of new technologies in electric furnace for

desulphurization, where used different reagents and then continued in LD converters during the oxygen refining

process [7]. The refining process in LD-oxygen converters consists of removing impurities from molten metal, relying

on their different affinity to oxygen. Most of the impurities are removed between gases or process slag. Sulphur and

phosphorus in ferronickel pass through the ore, reducer and fuel. Sulphur, which is organically bound to the charge

components, part of it is expected to pass through the gases during the pre-reduction process, part pass through the

electric furnace, where during the melting process approximately 50% is expected to pass through the gas and and the

remainder pass into metal and slag [4, 11].

From the GIBSS energy calculations, is confirmed the solubility of (S) to (Fe), as a result of which the molecule of

(FeS) is formed. The formation reaction of (FeS) is followed by large heat release effects [4], Steeluniversity.org Basic

2006 [12]. However, during blowing to the surface of the metal, a reaction may occur between the (S) and the (O2)

oxygen from the gases, which resulting in the formation of SO2 or SO3 which leave with gases. The gas created is

absorbed by the ventilator and it is likely that there will be no reversible reaction, NewCo Ferronickel [5].

The dynamics of removal of harmful impurities, depending on their concentration in ferronickel and oxygen flow

(Nm3/min). Reaction of sulphur metal with oxygen, under refining conditions in converter, is not possible [4, 13], no

matter that Gipss' free energy for this reaction is identical to Equation 1:

G0 = 1340 + 12.8T (1)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

911

The transition of sulfur from metal to slag has been the subject of many researchers. According to data from the

Cоколов (1982) research [14], the isotope of S35 and Fe59 are the first to transition from metal to slag and enable

reactions in the metal-slag system. Such a process in the metal-slag system, during refining in metallurgical reactors, is

enabled by the presence of basic oxides (MeO) and formation of (MeS) according to the basic reactions (2, 3 and 19):

(S) + (O2-

) = (S2-

) + (O) (2)

(FeS) + (S) = (FeS) (3)

After those, metal sulphide dissolved in metal can react with the calcium and nickel oxides in slag, according to

reactions 4 and 5:

(FeS) + (CaO) = (CaS) + (FeO) (4)

3(NiO) + 3[FeS] + O2 = (Ni3S2) + 3(FeO) + SO2 (5)

(FeS) = (FeS) (6)

Whereas;

(CaO), (FeO), (CaS) dhe (MeS) are components dispersed in slag and

(S) and (O) are components dispersed in metal.

Theoretically, reaction 4 and 6 begin to develop prior to the process of refining the converter, but reaction 6 may

continue to develop during the blowing with oxygen. During oxygen blowing conditions should be created for the

change of the standard isobaric potential for the oxidation reactions of Fe and FeS to occur. Sulphur activity in iron is

determined on the basis of concentrations of (S) and other elements involved in ferronickel melting Brankovič et al.,

1980 [15]. The desulphurizing capacity, which is determined by the sulfur separation coefficient, is depend on the

oxygen activity participating in the reaction, the metal composition, the sulfur to metal concentration, the reactivity of

the limestone used as defroster, process temperature, partial pressure of oxygen and sulphur in gaseous phase in

equilibrium with slag, chemical-physical composition of the slag and above all by; viscosity, surface tension, ratio

between oxides of SiO2, CaO, MgO, FeO, as well as other technological parameters that would stimulate the

interactive process in system metal-slag.

..

(S) +

(MeO) + Feliquid = (MeS) + (FeO)

(7)

Desulphurization mechanism is believed to consist of transfer of ionic iron pairs from metal to slag (reciprocally)

according to the scheme:

(S) + Feliq.. ↔ S2- + Fe2+

(8)

Equilibrium constant of reaction is usually described by equation:

𝐾𝑆 =𝑎(𝐶𝑎𝑆) ∙ 𝑎(𝐹𝑒𝑂)

[𝑆] ∙ 𝑎(𝐶𝑎𝑂) (9)

To increase the desulphurization rate, operating regime temperatures must be maintained at optimum values (optimum

values considered slightly above the melting temperature level of electrical furnace ferronickel), but maintaining at

optimum ratio (CaO)+(MgO)/SiO2 parameters. In acidic slurries the rate of ferronickel desulphurization is almost

negligible. According to B.P. Onshine, the desulphurization rate in the acidic medium reaches a maximum value of up

to 15%, shown in Figure 2 [4].

Slag desulphurization capacity towards the metallic phase may be expressed as the slag sulphide capacity:

𝐶𝑠 = (𝑠)

[𝑆]= 𝑓 [

𝐶𝑎𝑂

𝑆𝑖𝑂2, 𝑇,

1

𝐹𝑒𝑂] (10)

Slag’s compounds with high iron content will decrease the sulfur activity. Increasing the amount of FeO (above the

concentration limit of 18% FeO), will adversely affect the desulphurization effectiveness of molten ferronickel.

Equation 10 shows that the following conditions must be met for successful desulphurization:

That sulfur has this strong bond, in the form of (mes),

That chemical component (mes) to have good solubility in slag,

That the desulphurizing ability of meo is high,

That the left side of the reaction 10 to shows good dispersive capabilities of S, from metal to slag, and

Civil Engineering Journal Vol. 6, No. 5, May, 2020

912

The ratio (S)/(S) to be linearly dependent on ratio (meo)/(feo).

Figure 2. Dependence of transition of: (a) sulfur on slag from basicity of slag and (b) concentration of total Fe

In the case of oxygen bowling on the surface of the molten ferronickel, will be suitable conditions are created to

oxidize the sulfur together with the iron, whereby the isobaric-isothermal reaction potential Equation 11 changes at

temperature 1600 0C, ΔG0 = -134000 J/mol. Thus it is possible that this reaction is thermodynamically dependent on

partial pressure of oxygen on molten metal and transfer / passage conditions of this mass of sulfur in surface melting

slag [10].

S +O2 = SO2 (11)

Slag desulphurization capacity against the gaseous or metallic phase may be expressed as the slag sulphide

capacity [13]:

System slag – gas:

(O2-) + ½S2 = (S2-) + ½O2

(12)

𝐶𝑆 = (%𝑆2−) ∙ 𝑝𝑂2

𝑝𝑆2⁄

12

= 𝐾0 ∙ 𝑎𝑂2− 𝑓𝑆2−⁄

(13)

System metal-slag:

(S) + (O2-) = (S2-) + (O)

(14)

𝐶′𝑆 = (%𝑆2− ) ∙

𝑎[𝑂]

𝑎[𝑆]

= 𝐾0

𝑎𝑂2−

𝑓𝑆2−

(15)

Sulphur partition between slag and metal may be expressed using the sulphur partition coefficient LS:

𝐿𝑆 =(%𝑆)

[%𝑆]= 𝐾0

𝑎𝑂

2− ∙𝑓[𝑆]

𝑎[𝑂]∙𝑓𝑆2− (16)

Where:

a(O) , a(S) - oxygen and sulphur activity in molten metal,

aO2-

, aS2 - oxygen and sulphur ions activity in slag,

(%S2-

) - sulphur weight % in slag,

pO2 , pS2 - partial pressure of oxygen and sulphur in gaseous phase in equilibrium with slag,

Ko- reactions equilibrium constant,

f[S], fS2-

- Henry’s activity coefficient of sulphur in metal and slag [7].

It is possible to determine LS from known thermodynamic data of oxygen and sulphur dissolution in molten iron

using CS. Sulphur partition coefficient (LS) is not only the function of slag composition, but depends also on the

oxygen activity in metal, i.e. also on the metallic phase composition [13]. The value of the sulfur partition coefficient

0.02

0.024

0.028

0.032

0.036

1.6 2 2.4 2.8 3.2 3.6 4 5

The

conce

ntr

atio

n o

f S

in F

e-N

i (

%)

(a) The basicity of slag CaO/SiO2

0.022

0.026

0.03

0.034

10 12 14 16 18 20 24 26

The

conce

ntr

atio

n o

f S

in F

e-N

i (%

) (b) Total Fe in Fe-Ni (%)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

913

in case of ferronickel slag usually increases with the increase in CaO activity, which represents the precise size of the

basic slag and in the case of a decrease in oxygen activity. The sulfur equilibrium in metal-slag system is also highly

dependent on the level of metal oxidation. The relationship between [S] in this system is expressed as; (S) ≈ 4 (O),

Stirling [9].

The higher the oxidation of the metal (O), the higher the residual sulfur concentration in the metal (S). To this

dependence on desulphurizing mechanism , some metallurgist have given a different interpretation, according to

which there is a sulfur volume-oxygen between metal and slag of oxidizing smelter, which oxidizes metal according to

the scheme; (S) + O2- ↔ S2- + (O), [4]. According to this equation the separation coefficient of sulfur will be;

𝐿𝑆 = 𝐾𝑆 ((𝑂2−)

[𝑂]) (17)

The relatively high desulphurization rate (up to 40%) of the converter bath can be explained by a two-phase process,

involving the slag and a phase of the oxidizing gas by reaction;

(S) + (CaO) + Feliq. → (CaS) + (FeO)

(18)

(CaS) + O2 → (CaO) + SO2

(19)

When sulfur is separated from slag, its equilibrium is shifted to slag-metal system, which contributes to the further

partition of sulfur from the metal to slag, ie. deepens desulphurization. This fact derives from the high activity of

sulfur in the case of low oxidation. The development of reactions however is actually conditioned by the effectiveness

of the removal of the slag after desulfurization. This process is essentially similar to diffusive oxidation of steel baths.

4. Results and Discussion

Based on the average chemical composition shown in Table 4, it is observed that the electric furnace metal from

"Ferronickel Foundry" in Drenas in its composition, besides the high Si concentrations, the other impurity with high

concentration is also S, which presents difficulties in the process of desulfurization.

At electric furnace ferronickel, sulfur is most commonly present in form of FeS, as it accounts for most of this

melting [4, 16]. Sulfur activity in liquid iron is determined based on the concentrations of [S] and other elements

involved in ferronickel melting. So the chemical composition of the ferronickel melting of the electric furnace is what

determines the bond between sulfur and the liquid iron. Slag desulphurization capacity (Cs) depends on the chemical

composition of the slag, that is, the ratio (CaO) + (MgO)/SiO2) as well as the process temperature. Increasing the

amount of FeO in the slag, in the first and second phase of blowing with O2, positively affects the partition of sulfur

from the molten ferronickel, but this can only be up to the concentration of 18% FeO.

Table 4. Average of chemical composition of ferronickel

As Fe Si Cr C P S Ni Co Cu

0.03 72.5-84 2.7- 4.0 0.25 - 0.76 0.3- 0.6 0.07-0.2 0.7- 2.0 11.22 -17 0.5-0.74 0.03-0.05

Figure 3. Dependence activity of coefficient S on the concentration of the constituent elements of ferronicke

C

Si

P

Cu

Mn S -0.2

0

0.2

0.4

0.6

0.8

0

1

2

3

4

5

6

0 2 4 6 8

log ɣs

Influence of ferronickel concentration constituents on activity coefficient of S (%)

Sulp

hid

cap

acit

y C

S

Civil Engineering Journal Vol. 6, No. 5, May, 2020

914

Figure 4. Dependence of sulfur partition coefficient on concentration of (FeO) (Initial slag 2.0)

Further increase of (FeO) in slag shows negative effects on desulphurization. The desulphurization rate increases if

the process takes place only slightly above the melting temperature of the ferronickel. In the first refining period the

removal rate of S is not large shown in Figure 3, so there are large oscillations with respect to the increase in S in the

slag. In this case, for a volume unit of oxygen, three times as much gas is formed and about twice as much heat is

released, compared to the last refining stage [4, 13]. All reactions of the first phase of refining are followed by high

heat release at this phase, to reduce the temperature and soften the acidic medium, it is necessary to use a large amount

of limestone.

Although chemical composition of metals gained from electric furnace shown in Table 5, for comparative period

November 1997 and January 2019, does not present any major change, as observed from Figure 2 and Table 4, quality

of refined ferronickel and the average effectiveness indicators, desulphurization has a high difference in their selves.

Table 5. Comparison of average converter efficiency indicators for November 1997 and January 2019

Nr. of

charge

Resource costs Metal from electric furnace and refined

Du

ra

tio

n o

f ch

arg

e (

h)

Blo

w e

fec. (m

in)

Yea

r w

hen

ch

arg

e i

ts

fin

ish

ed

Analysis of molten ferronickel

Ca

CO

3 (k

g)

O2 (

Nm

3)

Gro

un

d m

ag

nesi

te

bric

ks

an

d C

aC

O3 r

ap

.

2:1

(k

g)

Fe-N

i fr

om

FE

(k

g)

S i

n F

e-N

i ra

fin

ed

(%

)

Scra

p (

kg

)

% N

i in

Fe-N

i

Si C S Ni

1719 6620 1043 - 19000 0.09 3000 43.19 161 55 1997 2.65 0.32 1.24 18.53

C-0831 5380 1855 360 22000 0.4 1500 18,06 121 56 2019 3.03 0.53 1.67 11.42

1722 8550 1214 - 18560 0.1 1500 42.43 180 68 1997 2.5 0.32 1.13 19.37

C-0833 6480 2101 670 19800 0.44 4500 17.69 111 58 2019 2.98 0.6 1.67 11.23

1724 7650 1169 - 20600 0.08 2600 45.88 153 63 1997 2.69 0.33 1.23 18.32

C-0835 6940 1743 840 21000 0.42 1500 16.72 73 53 2019 3.07 0.6 1.68 11.48

1726 9010 1414 - 18960 0.09 2500 45.95 180 50 1997 2.05 0.32 1.38 18.12

C-8441 5740 1942 600 23000 0.44 2500 17.61 112 55 2019 2.86 0.6 1.62 11.22

Civil Engineering Journal Vol. 6, No. 5, May, 2020

915

Figure 5. Reaction rates for oxidation reaction in BOS [5]

According to the process indicators, at the first refining phase, the desulphurization was very low, while at the

second phase when the composition of FeO in the slag it's about 18% FeO, then conditions are created favorable

conditions for desulfurization. Slag of second phase are characterized by low percentage of silicon oxide while high

percentage of CaO and FeO, shown in Figure 6.

Theoretically, if desulphurization was carried out of furnace using desulphurizer Mg, Na2CO3, synthetic slag or any

similar desulphurizer the desulphurization rate during this intermediate phase would have to be 40-60% [S], removed

from the electric furnace of ferronickel [3, 4]. Such desulfurization practices would stimulate the relationship between

the LS coefficient and the no. 2-/(O) ratio, favoring the formation of basic slag’s. In cases where favorable kinetic

conditions are created it is possible to provide additional desulphurization of the molten bath as the result of the metal-

slag reactions. During additional desulfurization an LS partition coefficient of 50 to 150 can be obtained Michalek,

2014 [10].

According to current parameters, virtually this mid-phase is eliminated, and all desulfurization is carried out on the

converter. This practice has resulted in high deviation of the sulfur partition coefficient from the metal between charge

1721 and C-0835, as well as the assumed coefficient of separation assumed to be developed according to theoretical

parameters shown in Figure 7 [7, 8]. The theoretical parameters of the partition constant were calculated assuming

magnesium desulfurization outside the furnace.

5 10 15 20 25 30 35 40 45 50 55 60

S/1721 1.18 0.897 0.625 0.42 0.216 0.17 0.142 0.116 0.086 0.073

S/C-0835 1.62 1.55 1.4 1.3 1.2 1 0.9 0.8 0.67 0.6 0.44

Si/1721&C/0835 1.5 0.5 0.3 0.1 0.03

C/C-0835 0.6 0.4 0.1 0.01

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0.5

0.55

0.6

0.65

0.7

0.75

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

1.2

1.25

1.3

1.35

1.4

1.45

1.5

1.55

1.6

1.65

Con

cen

tra

tion

of

S, S

i an

d C

for c

harge n

0. 1721 a

nd

C/0

835 (

%)

Effective refining-flow time (m3n/min) O2

S/1721

S/C-0835

Si/1721&C/0835

C/C-0835

Civil Engineering Journal Vol. 6, No. 5, May, 2020

916

Figure 6. Chemical composition of slag by process periods

Figure 7. Avoidance of sulphur partition coefficients for charge 1721 and C-0835 by partition coefficient according to

theoretical parameters out furnace desulfurization with Mg

The equilibrium state between the metal and slag would be reached quickly, that is, the desulphurization time

would be shortened, and action of reactive components between them would be at a high level if desulphurization

would take place outside the furnace, but above it, is necessary to precisely determine the boundary between the liquid

and solid temperature molten ferronickel, the partial oxygen pressure, presence of ferrous oxides in slag and degree of

slag basicity [17].

0

5

10

15

20

25

30

35

40

45

50

55

60

Con

cen

trati

on

of

slag b

y p

rocess

perio

ds

in (

%)

Formative slag attachments

zgj.e per.I

zgj.e per.ll

zgj.e per.lll

Civil Engineering Journal Vol. 6, No. 5, May, 2020

917

5. Conclusion

Considering the desulphurization of electric furnace ferronickel, which is produced at the “Ferronickel Foundry” in

Drenas, it is very important to discuss in the future the quality of Kosovo's lignite, which is used as a reducer. Such

lignite qualities have, as a consequence, created carbon-poor and highly sulphur-rich metal. The effectiveness of

desulphurization, optimization and economic effects process will be achieved through new desulphurization practice

application outside of furnace that would utilize desulphurizers, such as; Mg, CaO, synthetic slag, or even mixture of

Na2CO3 and FeSO4, etc. Practically in recent times the removal rate of sulphur from the furnace ferronickel has been

completely developed in the converter, by this process flaw, refined metal has been followed with concentration up to

0.45% S, production capacity level has dropped in parameters minimum utilization, normative costs of energy

resources and raw materials and all other circumstances to economize the production process have been unfavorable.

From the data of this study, it has been concluded that in addition to the application of mid-phase desulfurization

outside the furnace, which is, in fact, the main component of the effectiveness of the ferronickel refining process, it is

also important maintaining of optimum composition parameters and extended sulphide capacity, metal composition

(especially the drop in optimum values for Si and C), operating temperature, granulometric composition and proper

melting reactivity, as well as maintaining the optimum operating parameters.

6. Acknowledgments

Authors thank their colleagues from NewCo Ferronikeli, who support us and greatly assisted during the research

work.

7. Conflicts of Interest

The authors declare no conflict of interest.

8. References

[1] Schemmel, Th., Schade, L., Kouzoupis, P., Beqiri, F., “Magnesia-carbon refractory lining for ferronickel converters-

optimization and lining improvement at NewCo Ferronikeli (Kosovo)”, The Thirteen International Ferroalloy Congress,

Efficient Technologies in Ferroalloy Industry, Almaty, Kazakhstan, 2013.

[2] Zhang, Lifeng, and Brian G. Thomas. “State of the Art in the Control of Inclusions during Steel Ingot Casting.” Metallurgical

and Materials Transactions B 37, no. 5 (October 2006): 733–761. doi:10.1007/s11663-006-0057-0.

[3] Geerdes, Maarten, Renard Chaigneau, and Ivan Kurunov. Modern Blast Furnace Ironmaking: An Introduction (2009, New

Edition 2015). IOS Press, 2015. doi: 10.3233/978-1-61499-499-2-i.

[4] Ibrahimi I., “Desuphuring of ferronickel outside its furnace – a possibility for intensification and optimizations of the process of

obtaing ferronickel”, The Thesis of Mastery, Faculty of Mining and Metallurgy in Mitrovica, UP Prishtina, 2009.

[5] Refining Process Technology Card - NewCo Ferronickel; C-0835 – 0901, Glogoc/Kosovo, 2019.

[6] Л. Ͷ. Πимeнов, B. Ͷ. Μихaҋӆов, “Πeppaбotka okͶcлehhыx hͶeлebыx pуд”, Mосвa “Metaӆургͷя”, 1987

[7] Jiří Bažan, Ján Kret, “Iron and Steelmaking” Academic materials for the Economics and Management of Industrial Systems

study programme at the Faculty of Metallurgy and Materials Engineering, VŠB - Technical University of Ostrava, 2015.

[8] Elkem a/s dhe Scandinavian Cancers, Progami investiv dhe propozimi i praktikës së desulfurimit në “Furrën elektroharkore për

rafinim” Shkritorja e Ferronikelit Gllogovc (1986).

[9] Stirling, D., “The Sulphur Problems: Cleaning Up Industrial Feed Stocks.” Chemistry Department. University of Glasgow,

Royal Society of Chemistery, Cambridge, UK, 2000.

[10] Karel Michalek, CSc., “Electrometallurgy and ferroalloys production”, Academic materials for the Metallurgy engineering

study programme at the Faculty of Metallurgy and Materials Engineering, VŠB – Technical University of Ostrava, 2014.

[11] Z. Slovič, “Termodinamički pristup desulfuraciji pri vanpeč obradi kiseonič konvertorskog čelika”- Doktorska disertacija”,

Beograd, 2013.

[12] User Manual, “Basic Oxygen Steelmaking Simulation User Manual”, The University of Liverpool, 2006. Available online:

http://34.214.117.115/content/html/por/BOS_UserGuide.pdf (accessed on 14 March, 2020).

[13] Kijac, J., and M. Borgon. "Desulphurization of Steel and Pig Iron." Metalurgija 47, no. 4 (2008). Available online:

https://hrcak.srce.hr/file/41149 (accessed on 14 April, 2020).

[14] Г. A. Cоколов, “Проиводство стали”, Metaӆургͷя, Mосвa 1982.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

918

[15] Marica Brankovic, Srdan Markovic: "Livene Legure Zelezo - Ugljenik" TMF: Faculty of Technology and Metallurgy,

Beograd (1980). pp. 1-17.

[16] Irving, W.R., “Continuous Casting of Steel”, The Institute of Materials, The University Press, Cambridge, London, 1993.

[17] Ghosh, Ahindra, and Amit Chatterjee. “Ironmaking and Steelmaking: Theory and Praktice”. PHI Learning Private Limited,

2008, 472 p.

Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

919

Development of Filters with Minimal Hydraulic Resistance for

Underground Water Intakes

A. A. Akulshin a, N. V. Bredikhina a*, An. A. Akulshin b, I. Y. Aksenteva c, N. P. Ermakova a

a Real Estate Management and Mining Department, Faculty of Architecture and Engineering, Southwest State University, Kursk, Russia.

b OOO Ekopromservis, 305029, Kursk, K. Marksa str., 47, Russia.

c Corvinus School of Economics, Corvinus University of Budapest, Budapest, Hungary.

Received 17 December 2019; Accepted 06 April 2020

Abstract

The development of modern structures of water wells filtering equipment with enhanced performance characteristics is a

vital task. The purpose of this work was to create filters for taking water from underground sources that have high

performance, long service life, quickly and economically replace or repair in case of performance loss. The selection of

the filter device must be made taking into account all the geological features of the aquifers, the performance

characteristics of the filter devices and the size of the future structure. Filter equipment designs for water intake wells

have been developed in this study. These filters have low hydraulic resistance, high performance and are easy to repair.

This article presents the dependency of flow inside the receiving part of the well, the dependence of filter resistance at

various forms of the cross section of the filter wire and the selected optimal section. The paper proposes a method for

selecting the optimal cross-section of the filter wire used in the manufacture of a water well filter. The proposed

structures of easy-to-remove well filters with increased productivity allow replacing the sealed well filter with a new one

easily, reducing capital and operating costs, and increasing the inter-repair periods of their operation. Based on the

presented method, examples are given for selecting the parameters of the filter wire cross-section. The above calculations

showed that the use of the hydraulic resistance criterion at the design stage of underground water intakes can

significantly reduce the cost of well construction. Studies have found that the minimum hydraulic resistance to ensure

maximum filter performance is achieved when using filter wire teardrop and elliptical shapes.

Keywords: Hydraulic Resistance; Filter; Water Well; Resistance; Filter Wire; Capacity.

1. Introduction

Groundwater use is 70% of the total water consumption in some European countries with the best rates in quality of

life for the public. These countries are Germany, Austria, Denmark, Belgium, Switzerland, and several others. More

than 300 million groundwater intake structures have been drilled around the world over the past 25-30 years [1-2].

The use of groundwater for water supply of the population has many significant advantages. High water quality in

the water supply source helps to avoid the necessity for preparation equipment using due to protection from external

contamination and seasonal changes of indicators. The capacity of a water well depends on hydrogeological

characteristics of soil and groundwater, as well as the structure of intake portion of a well and pumping equipment.

Therefore, the development of modern structures of water wells filtering equipment with enhanced performance

characteristics is a vital task.

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091517

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

920

The main element of a downhole water intake is a water intake well. The quality of its design and construction

influences the water intake operation. The main requirements for a water intake well (tubular well) are as follows:

production of the required amount of water with the quality corresponding to the requirements of consumers,

efficiency and reliability in operation. The main elements of the well design are: guide column, conductor,

intermediate columns (technical columns of casing pipes), production column, cement or other protection and water

intake part. The purpose of tubular well filters is to keep the soil from collapsing and at the same time ensure free

passage of water into the well bore. Filters must meet the requirements:

Have the necessary mechanical strength;

Ensure that water flows into the well with minimal hydraulic resistances and without mechanical impurities;

Have high corrosion resistance;

Ensure maintainability and the ability to extract from the well.

The following types of downhole filters are used: frame-and-rod; tubular with slotted or round holes; with wire

winding; reticular; polymer ring. Their choice depends on hydrogeological conditions, depth of occurrence and types

of rocks of aquifers. Frame-and-rod filters are considered to be the most efficient one. A single filter surface (profiled

wire) is available for chemical and mechanical cleaning, since there are no dead spaces between the filter and support

surfaces. The consumption of metal on a core frame is about half that of a perforated pipe frame. Filters on core

frames are used in wells up to 200 m deep.

Filters on a tubular frame with a wire winding are also common due to the ease of manufacture. The duty cycle of

these filters is 20-25%. Their disadvantage is that there is an accumulation of colmating connections between the wire

and the frame. The cross-section shape of the filter wire can be round, teardrop, elliptical, trapezoidal, rectangular, or

other. Wedge Wire Screens made by “Johnson”, pressed filters with tunnel-holes and plastic slot-type filters are

popular abroad [3-5].

The most common filters in Russia are mesh and wire filters with perforated tubular frame. The main reasons that

impede the selection of a filter are the difficulty of determining the filter's open ratio and the variability of the distance

between the loops of the filter’s round wire during manufacture. In our opinion, structures of easy-removable frame-

and-rod filter-frames which have a fixed size hole and made of anticorrosion materials are advanced ones. However,

water intakes from underground aquifers also have drawbacks. The most common problem of well operation is

incrustation. Regular well cleanout has a temporary effect [6-8]. The well capacity gradually decreases with increasing

resistance to the flow of water from an aquifer. This demonstrates the importance of the development of construction

technology of wells that have stable debit and can withstand the harmful effects of fine sand-grains contained in the

pumped water, which leads to incrustation and severe abrasion of structural elements of an underwater pump [9].

2. Development of Easy-Removable Filter Frames with a Fixed Size Hole

The tasks of designing well filters are to determine the optimal design and technological parameters. The main

technological parameters of borehole filters are their borehole capacity and hydraulic resistances of the filter surfaces.

Increasing the filter life allows you to reduce its length and diameter, and therefore the cost of filters.

Figure 1. The nature of filtration in the well, taking into account the minimum energy costs

Civil Engineering Journal Vol. 6, No. 5, May, 2020

921

The length of the filter determines the capacity, structure, and type of aquifer. The length of the filter exceeding 10

m is not appropriate; this is due to the fact that the length of the filter reduces the load on it from the pump unit to the

sump located in the lower part of the well bore. Calculations carried out by V. G. Tesla show that when the filter

diameter increases from 168 mm to 325 mm, the specific flow rate increases only for 9.2%, but the cost of

constructing with a larger diameter well may increase more than twice [7].

The physical meaning of the flow pattern easily explained in Figure 1. By integrating the pressure plot according to

the reservoir power, you can determine a point in the reservoir at an equal distance from the well, to which the

pressure gradient of the moving flow is directed in any interval. The location of the desired point will be significantly

shifted from the upper limit of the exploited interval. Therefore, it is advisable for the flow in remote areas

characterized by low velocity (v) to move in the direction of the pressure gradient, i.e. in a radially spherical flow. As

the flow moves toward the well, its cross section decreases and the flow rates increase significantly, which means that

the hydraulic head loss increases. At a certain distance from the well, the flow begins to run out, the "live" section

increases, which helps to reduce the filtration rates and transition to a more energy-efficient form of movement.

Despite the fact that the length of the current line increases, the possible increase in pressure losses due to this is

compensated by their decrease due to a decrease in filtration rates.

Taking into account one of the basic laws of hydraulics, which assume the flow movement along the path of least

resistance with minimal energy costs, a radial - spherical flow is formed in the reservoir in remote areas, which begins

to transit gradually to a flat-radial one in some areas. The highest flow rate is observed in the upper intervals of the

reservoir, where the current line thickens to the maximum. In the lower intervals of the reservoir, the frequency of the

current line is significantly reduced due to the mismatch of the direction of movement with the pressure gradient,

which indicates lower inflow intensity than in the upper intervals.

One of the most important conclusions that follow from the presented diagram is the possibility of determining of

the formation part that is intensively loaded. The highest load is taken by the upper intervals of the formation, which

are separated from the upper boundary at a distance of 𝑚 = 𝑟кр𝑚1/𝑅. Well performance is determined by hydraulic

head losses in all sections of the flow. The coefficients of laminar and turbulent hydraulic friction are generalized.

There is a relationship between the well performance and the resulting drop in levels corresponding to the head loss in

the well-formation system. The resistance coefficients are generalized and are considered as a function of the sum of

the laminar and turbulent resistances of each element of the system. The most optimal operating modes should

provide a laminar filtration mode in all elements of the hydraulic system of the water intake well.

In the case of laminar filtration mode, an increase in the reduction is accompanied by a directly proportional

increase in well productivity. When the flow is turbulized on one or more traffic elements the decrease begins to be

accompanied by an increasingly slow growth of productivity. If the system is switched to a turbulent operation mode,

the decrease does not lead to a significant increase in the flow rate and operation becomes economically unprofitable.

The relationship between well performance and reduction is permanent only in the case of steady-state operation. In

real conditions, at the initial moment of operation or testing by pumping, developing levels, connecting and shutting

down neighboring wells, and others, the dependence of the reduction and flow rate begins to change over time. This is

due to the inertia of the well-formation system. It should be noted that with an increase in productivity caused by a

decrease in hydraulic resistance in the elements of the well - formation system, the head distribution will change step

by step. Increasing well productivity can be achieved by reducing the hydraulic resistance to flow movement on one

of the elements of the well-formation system.

Scientific publications and patents research of the existing design and technological solutions for the water wells

incrustation prevention provides the basis for the following conclusion: one of the most effective solutions to this

problem is to ensure easy replacement of the incrusted well filter during operation. Several core frame filters with

various forms of wire were developed in the South-West State University together with the company

"Ecopromservice" [10-13]. The Figure 2 shows a diagram of a core frame well filter [14-16].

The filter has the following construction [17-19]. The filter support frame (1) is a perforated metal pipe. It is a

continuation of the drive tube (6) lower part [20]. The support frame is designed to hold gravel pack in the aquifer and

to fix the filter units in the face of the tubular well. The support frame is located outside the concertina wire (2) that is

applied to the sag rods (3). There are vertically fixed centring slabs (7) with the angle 120 relative to each other

between them. It helps to avoid the filter shifting towards the well center. The centring slabs are anchored on the

connections sleeves (4, 5). Sag rods has ellipse configuration and are located outside the connections sleeves. The well

filter height can be adjusted by adding filter sections. The filter can be removed by a gripping tool in case of repair or

replacement, which is fixed to the lower surface of the upper coupling (8) and uses a traction device to remove the

filter from the borehole.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

922

Figure 2. Structure of the tubular well filter: (1) Support frame, (2) Concertina wire, (3) Sag rods, (4) Upper connection

sleeve, (5) Lower connection sleeve, (6) Drive tube, (7) Centring slabs, (8) Lower part of upper connections sleeve

3. Hydraulic Modeling of the Intake Part of the Well Filter

Geological features of an aquifer, performance characteristics of filtering devices and the size of the future structure

must be considered for the filtering device selection. Besides, it is important to have minimal hydraulic resistances for

water entry into the filter elements. The lowest hydraulic resistances do not depend on hydrogeological conditions and

are provided by the use of filters with slits oriented in the horizontal plane [21-23]. The most suitable filters that meet

this condition are wire-wound frame-rod filters. The performance characteristics of a well are primarily determined by

the presence and the pressure loss rate of intake portion that depends on the filter structure, drilling-in methods and

other factors [24].

Previous studies have shown that there is an optimal number of holes (optimal borehole). The increase of the

number of holes, although it contributes to the overall increase in the total flow rate, leads to the reduction of

resistance and slowes down due to the increased interaction of holes (interference effect). Thus the more holes will be

put into operation the overall effect of increasing of the total debit will decrease faster [25-27].

Let's consider some hydrodynamic solutions to the problem of inflow to a well equipped with a filter, without

taking into account the imposition of rocks and reducing the well life of filters, as well as the possibility of chemical

overgrowth of filters. At the same time, filter resistance can be evaluated only in conditions of captage of stable rocks,

when the effect of overlaying water-bearing rocks on the water-receiving part of wells does not affect. In some cases,

the values of filter resistances will be underestimated. Since the geometry of the filtration flow near the well is

determined by the shape of the inlet holes, it is best to classify filters for these purposes on this basis. Analytical

solutions for the flow of liquid to a filter with round holes were obtained by M. Masket and A. L. Hein. When solving

the problem, round holes were replaced by drains placed along the filter pipes. It is obvious that such a scheme does

not provide physical similarity, since the impenetrability of the walls is not taken into account, but it is assumed that

this disadvantage is compensated by the effect of effluent interference [25-27].

It was found that the value of the filter resistance is almost independent of the well diameter. V. I. Shchurov

compiled curves by approximating M. Masket's analytical solutions using the same type of empirical equations for

three hole diameters: 6.4, 12.7, and 19 mm. The value of the filter resistance depending on the number of holes and

their diameter for the filter installed in a homogeneous formation. These dependencies were studied in detail By V. I.

Shchurov using the method of electrohyrodynamic analogies and as a result, refined graphs of the filter resistance

dependence on the parameters 𝛼 = 𝑑0/𝐷 and 𝛽 = 𝑛𝐷 (𝑑0 is the diameter of the holes, n is the number of holes per 1m

of the filter length, D is the filter diameter) were obtained [25-27].

Civil Engineering Journal Vol. 6, No. 5, May, 2020

923

A. L. Hein obtained solutions for determining the inflow to a filter with round holes in conditions of unsteady

movement. As a result, it was proved that the effect of unsteady flow in the filter zone can be traced for very short

period, so in practical calculations of filters, we can limit ourselves to considering the stationary filtration mode [25-

27].

Filters with vertical slits are divided into two groups: filters with slits whose length is equal to the thickness of the

reservoir, and filters with slits of limited length. The first group includes filters made of rods without wire winding.

When determining the resistance value for these filters, very similar results were obtained using various methods.

These results were verified by the experimental method of electrohyrodynamic analogies, and satisfactory results were

obtained. A similar result was obtained by A. L. Hein when calculating the resistance of a filter with vertically

positioned holes without limiting the borehole. To determine the resistance of a slotted filter with rectangular slits of

limited length placed vertically along the filter formation, A. L. Hein obtained a corresponding solution, and on the

drainage surface, he assumed a certain average velocity of the liquid movement, and on the impervious sections of the

filter – a potential gradient equal to zero. This solution is very cumbersome and does not lead to calculated

dependencies. A. L. Hein's calculations show that the approximate resistance of filters with vertical slits of limited

length can be found by graphs of V.I. Schurov for filters with round holes by reducing a rectangular hole to an equally

large round one [25-27].

In this paper, we consider the hydraulic resistance of the filter without taking into account the contact head loss.

This resistance depends on the shape of the holes, their number (borehole), the size that determines the flow

dispersion, the nature of the location on the water-receiving surface of the well, and their interaction. The filter

hydraulic resistance is the most important part of the total resistance of the near-wellbore. The distribution of head

losses, intake velocity, and water influx rate should be considered for identifying common factors of flow movement

within the intake portion of a well. The following conditions were accepted for a hydraulic model of the intake portion

of a well development. The filtering surface consists of wire loops on the frame. It is considered as a screen area that

has equal thickness of the rods and the wires. The width of the rods gaps is equal to the spacing between the wire

loops.

The liquid that flows towards filtering surface is compressed in its holes and leave these holes as separate streams

with a high speed for the face of a well. Thus, these losses are associated with water inlet and sharp expansion at the

filtering surface outlet into the filter internal part (Figure 3). The coefficient of resistance of the filter surface depends

on the coefficient of the live section, the shape of the edges of the holes and the Reynolds number. At low coefficients

of the live cross-section of the filter surface, the flow rate in the holes can reach high values, especially in places

where the streams are most compressed. If the cross-section speeds are not evenly distributed, the filter surface aligns

the incoming flow. The resistance created by the filter surface redistributes the incoming liquid flow over the surface

and at the same time allows the liquid to pass through the filter holes.

The degree of alignment of the liquid flow on the filter surface depends on its geometric parameters. Since these

parameters determine the coefficient of resistance of the filter surface, the results of liquid flow redistribution are a

function of the coefficient of hydraulic resistance. As the coefficient of hydraulic resistance increases, the degree of

flow alignment over the area of the filter surface also increases. However, thin-walled filter surfaces, unlike bulk

obstacles, have their own characteristics: when a certain value of the hydraulic resistance coefficient is reached at the

outlet of the filter surface, the velocity profile becomes inverted in the opposite direction. At the same time, it can be

observed that the flow is uneven, where the maximum speed behind the filter surface corresponds to the minimum

speed in front of it [25-26]. When the liquid flows along the front surface of the filter, the flows are curved, since the

filter is thin-walled, the holes have no guide surfaces, and the transverse direction of the flow is maintained after the

liquid flows through the filter surface. This leads to further spreading of the liquid and movement of the flow in the

radial direction. The greater the hydraulic resistance of the filter surface is, the sharper the flow curve becomes and

there is a significant curvature of the flows coming out of the filter holes [25-27].

Figure 3. Water flows through the filtering surface

Civil Engineering Journal Vol. 6, No. 5, May, 2020

924

Two filtering surfaces set-up close to each other should not lead to an increase in resistance because the fine

alignment of both surfaces is the reason of the holes increasing along the stream. The filtering surfaces partially

overlap each other. That is why the flow section decreases slightly and the resistance increases. The total resistance

can be defined as the sum of the resistance coefficients of individual parts in case of two filters set-up at some distance

[25-26]. In that case, the distance between the sag rods, on which the filtration wire is wound, is much larger than the

diameter of the wire. That is why compensating resistance can be neglected. It is known that the overall hydraulic

resistance of any element of a chain is determined by the Equation 1 [26]:

∆𝑝 = 𝜉𝜌𝜔1

2

2= 𝜉

𝜌

2(

𝑄

𝐹)

2

(1)

Where; ∆𝑝 – Stagnation pressure reduction (Pa); ξ – Hydraulic resistance coefficient; ρ – Liquid density, kg/m3; 𝜔1 –

flow velocity (m/sec); Q – volume fluid flow rate (m3/sec); F – total area of filtering holes (m2).

The overall losses of filtering surfaces that are made of various form of section wire consist of inlet losses, friction

losses, and sharp expansion loses at the outlet. Some form of section and filter wire parameters are shown in Figure 4.

1 2 3

Figure 4. Form of section of filtering wire with minimal hydraulic resistance: (1) Drop-shaped; (2) Elliptic; (3) Circular

If l/dm = 5 and d0/S1 ≥ 0.5, coefficient of screen resistance can be determined by Kirshmer’s Equation [26-28]:

𝜉 =∆𝑝

𝜌𝜔12

2

= 𝛽1𝑘1 sin Θ (2)

Where: 𝛽1 – coefficient of rods shape (Table 1); 𝑘1 = (𝑆1

𝑑0− 1)

43⁄

(Table 2, Figure 5); 𝑘1 = 𝑓 (𝑑0

𝑆1); θ – Angel of wire

slope to flow; 𝑑𝑚 – Width (diameter) midsection of filter wire (m); 𝑑0– space between two adjacent wire loops (m); S1

– Space between two adjacent wire loops axis (m); l – Sectional length of filtration wire (m).

Table 1. Coefficient value β1 for various rods shape

rods 1 2 3

β1 0.87 0.71 1.73

Table 2. Value 𝒌𝟏 = 𝒇 (𝒅𝟎

𝑺𝟏)

𝒅𝟎 𝑺𝟏⁄ 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

k1 ∞ 18.7 6.35 3.09 1.72 1.00 0.58 0.32 0.16 0.05 0

Civil Engineering Journal Vol. 6, No. 5, May, 2020

925

Figure 5. Dependency graph

According to Equation 1, it follows that the volume fluid flow rate is equal to:

𝑄 = 𝐹√2∆𝑝

𝜉𝜌= 𝐹√

2∆𝑝

𝛽1𝑘1 sin Θ 𝜌 (3)

Thus, the volume fluid flow rate is inversely related to the hydraulic resistance. Accordingly, the greater the value of

hydraulic resistance the lower the filter capacity.

As a matter of practice, the filtration wire of circular section is commonly used. However, according to the Table 1,

drop-shaped and elliptical wires have the smallest value of hydraulic resistance. Filter capacities that have circular

section wire 𝑄𝑤𝑖𝑟, elliptic 𝑄𝑒𝑙 and drop-shaped 𝑄𝑑−𝑠ℎ𝑎𝑝 were compared. If 𝛽𝑤𝑖𝑟 = 1.73, 𝛽𝑒𝑙 = 0.71, 𝛽𝑑−𝑠ℎ𝑎𝑝 = 0.87

and other equal terms:

𝑄𝑤𝑖𝑟 = 𝐹√2∆𝑝

𝛽𝑤𝑖𝑟𝑘1 sin Θ 𝜌= 𝐹√

2∆𝑝

1.73𝑘1 sin Θ 𝜌= 1.08𝐹√

2∆𝑝

𝑘1 sin Θ 𝜌 (4)

𝑄𝑒𝑙 = 𝐹√2∆𝑝

𝛽𝑒𝑙𝑘1 sin Θ 𝜌= 𝐹√

2∆𝑝

0.71𝑘1 sin Θ 𝜌= 1.7𝐹√

2∆𝑝

𝑘1 sin Θ 𝜌 (5)

𝑄𝑑−𝑠ℎ𝑎𝑝 = 𝐹√2∆𝑝

𝛽𝑑−𝑠ℎ𝑎𝑝𝑘1 sin Θ 𝜌= 𝐹√

2∆𝑝

0.87𝑘1 sin Θ 𝜌= 1.52𝐹√

2∆𝑝

𝑘1 sin Θ 𝜌 (6)

After transformations Equation 4 to 6, the following results were obtained:

𝑄𝑑−𝑠ℎ𝑎𝑝 =1.52

1.08𝑄𝑤𝑖𝑟 = 1.41𝑄𝑤𝑖𝑟 (7)

𝑄𝑤𝑖𝑟 =1.7

1.08𝑄𝑤𝑖𝑟 = 1.41𝑄𝑤𝑖𝑟 (8)

𝑄𝑒𝑙 =1.7

1.52𝑄𝑑−𝑠ℎ𝑎𝑝 = 1.41𝑄𝑑−𝑠ℎ𝑎𝑝 (9)

Equations 7 and 8 shows that the capacity of filters with drop-shaped wire is 1.41 times more than the capacity of

filters with circular wire and 1.57 times more than the capacity of filters with elliptic wire. In our opinion, the most

promising designs of downhole filters are frame-rod, easily removable filters with a fixed slot size and using a drop-

shaped filter wire made of materials that are resistant to corrosion.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

926

4. Conclusion

Water wells with incrusted filters research have shown that the best solution for lifetime extension is to use in-place

repairable, highly efficient and easy-removable filters. The main requirements for a water intake well (tubular well)

are the extraction of the required amount of water with a quality that meets the requirements of consumers, as well as

cost-effectiveness and reliability in operation. Important technological parameters of downhole filters are their

downhole capacity and hydraulic resistances of filter surfaces. Increasing the working cycle of the filter reduces its

length and diameter. The length of the filter determines the capacity, structure, and type of aquifer. Filter length

exceeding 10 m is not suitable.

The selection of a filter device must take into account all the geological features of aquifers, the operational

characteristics of the filter devices and the size of the future structure. At the same time, it is necessary to ensure that

the hydraulic resistance for entering incoming water into the filter elements is minimal. The use of filtering wire with

the drop-shaped form of section allows reducing hydraulic losses during fluid flux through filters. Therefore, it

improves the performance characteristics of water wells and minimizes capital commitment by extending the overhaul

periods. Studies have shown that the minimum hydraulic resistance to ensure maximum filter performance is achieved

by using a filter wire with a teardrop and elliptical shape. In our opinion, the most promising designs of downhole

filters are frame-rod, easy-to-remove filters with a fixed slot size and using a teardrop-shaped filter wire made of

materials that are resistant to corrosion. The resistance created by the filter surface redistributes the incoming liquid

flow over the surface. The degree of alignment of the liquid flow on the filter surface depends on its geometric

parameters. As the coefficient of hydraulic resistance increases, the degree of flow alignment over the area of the filter

surface also increases. The designs of easily extracted downhole filters of increased productivity are proposed to

eliminate the possibility of filter breakage during production, i.e. to provide easy replacement of a sealed downhole

filter with a new one.

5. Conflicts of Interest

The authors declare no conflict of interest.

6. References

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Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

928

Ranking and Determining the Factors Affecting the Road

Freight Accidents Model

Masoud Bagheri Ramiani a*

, Gholamreza Shirazian b

a PhD Student of Transportation, Department of Civil Engineering, Shomal University, Amol, Iran.

b PhD of Transportation Engineering, Department of Civil Engineering, Shomal University, Amol, Iran.

Received 15 November 2019; Accepted 04 March 2020

Abstract

The tremendous growth of population, particularly in developing countries, has led to increased number of travels,

especially those with load and freight specifications. Hence, expanding the present facilities or developing new networks

or systems concerning freight and transportation is an essential issue. Among the various transportation systems, road

freight has secured a significant place in sub-urban transportation, as it is responsible for transporting loads, decreasing

transportation costs, and increasing the safety of highway users. Besides these advantages, poor and nonstandard design

and performance of sub-urban highways and transport fleet and equipment leads to the increased number of accidents

and inefficiency of these facilities. Based on these facts, the primary aim of the present study is to probe into the factors

affecting road freight accident severity. For this purpose, the data obtained from road freight accidents occurring in 2016,

2017, and 2018 in Gilan Province, Iran, were used for analyzing the frequency, ranking and determining the factors, and

creating models for accident severity. The results indicated that in accordance with the accident severity model in 2016,

several factors such as the season of autumn, daytime light, drivers aged from 18 to 60, and pickup trucks have impacted

the on-road freight accident severity. While, in 2017 the severity was affected by factors like rural road, freight trucks,

non-faulty passenger cars, motorcycles, and pedestrians. When considering the effective variables in 2018, it was found

that such factors as the accident time (usually occurring between 12 p.m. to 6 p.m)., rural and major roads, freight trucks,

non-faulty motorcycles, and the careless driving without due care and attention to the front were the variables affecting

road freight accidents. Moreover, not following safety guidelines during freighting is the most effective variable in road

freight accidents.

Keywords: Accidents; Road Freight; Cargo; Damage; Injury; Fatality.

1. Introduction

Over the last decade and also nowadays, road accidents have experienced an increasing trend; and there have been

various studies on accidents around the world [1-3]. Road accidents are considered to be a typical phenomenon all

over the world and approximately 1.3 million civilians die as a result of this phenomenon. Moreover, approximately

20 to 50 million people have been injured in these accidents, where the majority were young people with ages ranging

from 15 to 45. Road accidents are claimed to be the ninth most important factor of fatality in the world as it accounts

for 2.2% of the mortality rate in the world. The costs of accidents are estimated to be about 500 million dollars all over

the world. This is equal to between 1 to 2 percent of GDP in the countries with low to average income. The current

trend of accidents indicates that if emergent measures are not taken in this regard, it is likely that road injuries will

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091518

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

929

become the seventh most important factor of fatality by 2030, 90% of which will have occurred in the countries with

low to average income. It is essential that emotional and mental injuries as well as permanent disabilities resulting

from these accidents be added to these complications [4].

Analyzing the accidents, it could be concluded that normally several parameters account for the accidents, which

can be categorized into human, environmental, and vehicle-dependent factors [5]. Various studies have revealed that

some parameters like AADT, traffic congestion, exceeding the speed limit and the number of lanes, drivers’

distraction, and weather conditions are among the factors which affect accidents [6-8]. Increase in the severity of any

of the above-mentioned factors and/ or any inappropriateness of these conditions can lead to increased number of

accidents [9-11].

An increase in the number of accidents results in the increase in the number of owners of motor vehicles. It can be

said that the mentioned growth has been about 65% over the last 2 decades; while in developing countries it has

occurred faster than this [12]. Because of this, road accidents are considered to be one of the most significant issues of

public health in any given society. Besides, it can be concluded that this problem is more serious than other public

health-related issues as the majority of its victims are young and healthy people [13, 14].

Ghaffar et al. (2004) [15] evaluated the effects of road traffic injuries (RTI) in Pakistan. The results indicated that

most accidents happened between 12 and 18 pm. The level of RTI is higher among people aged between 16 and 45

years. Furthermore, the results showed that RTI is almost three times higher in males than females.

Labinjo et al. (2009) [16] provided a population-based survey to explore the epidemiology of RTI in Nigeria. The

results showed that motorcycle accidents accounted for 54.33% of all RTI. The risk of crashes was higher among

males aged between 18 and 44.

Hu et al. (2012) [17] explored the characteristics of traffic accidents on rural roads using the quantitative analysis.

The research shows that 92.68% and 5.42% of casualties occurred on tangents and curves, respectively. Considering

the time parameter, casualties during the daytime have been more serious than the night. Also, the results showed that

crashes that cause injuries are most common during the day rather than night and motor vehicle accidents account for

the majority of casualties. Zangooei Dovom et al. (2013) [18] explored the fatal accident distribution in Mashhad, Iran.

According to the results, the male had more fatalities than the females. For both genders, most accidents had a peak at

the ages of 21-30. The male to female overall casualty ratio was 3.41. Among all the road users, the riskiest group was

male motorcyclists.

Lee and Jeong (2016) [19] investigated the characteristics of traffic collisions occurred in expressways and rural

roads among the truck drivers. The results showed that with respect to the day of the week, the accident rate was

higher in the middle of the week. On rural roads, the accident rate during the daytime was much higher than the night

time (81.7%). The accidents occurred mostly in clear/cloudy weather (76.2%). Besides, the majority of accidents

occurred over a straight road (62.2%), followed by an intersection (15.4%) and a curved road (9.4%).

Road freight is the oldest way of transporting cargo in the world. In terms of the price and speed, it is the most

appropriate way to transport a variety of goods and cargo. Nowadays, the majority of the cargo is transported by

means of road freight, which along with its merits, has some disadvantages. One of the disadvantages is the accidents

occurring due to the failure of freight vehicles on roads, which in addition to economic losses results in fatality or

injury of road users as well. Hence, in this study, the effective factors are identified by means of considering the

severity of road freight accidents, particularly the accidents resulting from some vehicles like pickup trucks, trucks,

trailer cars, etc. Then, through modelling and statistical analyses of accidental data, the appropriate approaches to

decrease road accidents and increase civil welfare and traveller’s safety are identified. For this purpose, after analyzing

the frequency of accidents, their ranking is performed and the factors affecting the severity of accidents are

determined. Moreover, the impact of independent variables on the severity of freight vehicle accidents is modelled.

The purposes and adopted innovations in the present study can be summarized as follows:

Application of Friedman test and Factor analysis methods for the road freight accidents,

Investigating the effect of independent variables on the severity of freight vehicles,

Frequency analysis of variables affecting freight vehicles accidents,

Ranking independent variables affecting the severity of freight vehicles,

Modelling of independent variables affecting the dependent variables of freight vehicles accidents severity,

2. Research Methodology

In this section, initially, the specifications of the study area will be introduced. Then, different statistical analysis

methods and modelling of accident severity will be executed according to Figure 1.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

930

2.1. Case study

Gilan is one of the northern provinces of Iran, whose capital is Rasht megacity. This province lies along the

Caspian Sea and Azerbaijan, sharing with it an international boarder via Astara in the north. It is located to the west of

Mazandaran Province, east of Ardabil Province and north of Zanjan and Qazvin Provinces. Gilan covers an area of

14044 square kilometres and based on the census carried out in 2012, its population is 2480874. Gilan is the tenth

province in Iran in terms of population and is the second most populated province in northern Iran i.e. it ranks only

second to Mazandaran Province. The population density of this province is 177 persons per square kilometre, which

secures third place in Iran. Constituting 46% of the total population of the province, Rasht megacity is the center of the

province and the most populated city in the north of the country and the 11th most populated city in Iran [20, 21].

Sub-urban highways of Gilan Province are 2573 kilometres in length, of which 1682 km, i.e. 65% of the total

length of highways within the province and 2.2% of the total length of highways of Iran, are capable to be utilized for

freight purposes. There are 363, 256, and 1063 kilometres of the mentioned total highway network length function as

highway, main, and rural roads, respectively. Figure 2 displays all the existing roads in Gilan Province, which can be

used for freight transport. Due to the abundance of details, the functions of rural roads are excluded [22].

Figure 1. Flowchart of the present study

Figure 2. Representing the roads capable of freighting in Gilan Province (Iran)

Accident data collection from 2016 to 2018

Identify freight accidents

Delete incorrect

information

Data

encoding Input to SPSS

Analysis and Modelling

- Frequency analyses

- Ranking by Friedman test

- Determining the factors by

factor analysis - Modeling accident severity

Conclusion

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2.2. Statistical Analysis and Modeling Methods

For analysing the Road Freight Accidents of Gilan Province, some statistical analysis and modelling methods,

including Kolmogorov–Smirnov test (K–S test), Friedman analysis, Factor analysis, and Logit modelling were used.

The statistical analyses used in this study were performed using SPSS software.

2.2.1. Kolmogorov-Smirnov Test (K–S test)

One of the main assumptions for most statistical tests is the normality of data distribution where the Kolmogorov–

Smirnov test (K–S test) is utilized for this purpose. This test is a nonparametric test for data distribution. In

approximate significance test, comparing | the output with α (significance level), the normality of data distribution can

be determined. If α =0.05(means with 95% certainty) if P-value >0.05, the data distribution can be assumed as normal.

Indeed this test is a compliance testing of the quantitative data distribution. Normality distribution test is the most

common test for examining the normality of a specific distribution [23].

2.2.2. Friedman Test

The Friedman test is one of the statistical tests used to compare between several groups and, ranks groups by using

the average value, whether these groups belong to one community or not. This test is a non-parametric one

corresponding to the F test and is usually used in ranking scales rather than the F test [24]. In the F test, there should

be homogeneity of variances that is less observed in ranking scales. The Friedman test is applied for the analysis of

two-way variance (for non-parametric data) by a ranking method. Also it is used to compare the average ranking of

different groups.

2.2.3. Factor Analysis

The factor analysis method is used to find out the underlying variables of a phenomenon or for summarizing a set

of data. The primary data for factor analysis is the matrix of correlation between variables. Factor analysis does not

have predetermined dependent variables. Factor analysis is applied for two general categories: exploratory purposes

and confirmatory purposes. If there is no speculation about the structure of the dimensions relationships, exploratory

factor analysis is used. Otherwise, the confirmatory factor analysis is used [25].

In the exploratory factor analysis, the researcher seeks to investigate the empirical data to discover and identify the

indices as well as the relationships between them. There is no pre-defined model here. In other words, exploratory

analysis, in addition to its exploratory or suggested value can be a structure maker, modeller, or hypothesis creator.

Exploratory factor analysis is used when the researcher does not have sufficient previous and pre-empirical evidence

to create a hypothesis about the number of underlying factors and wants to determine the number or nature of the

factors justifying the covariance of variables. Therefore, exploratory analysis is more considered as a method of

compilation and production of a theory, rather than a method of testing a.

2.2.4. Multiple Logit Regression

Establishing a relationship between the set of variables x and the dependent variable Y, we would encounter a

multivariable problem. In analysing such a problem, various types of mathematical models have been used to consider

the complexity of the relationship between these variables. The logit regression method is a mathematical method used

to describe the relationship between multiple variables denoted by x and a two-valued dependent variable. A function

that is used in this method is an S-shaped function called the logit function, which can also be applied in multi-valued

problems by expansion [26]. As it is known, the logit regression method can be utilized to define the variable Y as the

multi-valued parameter. In the simplest case, we can consider P(Y=i) as a linear function of XI (Pi = xi β), where β is

the vector of regression coefficients. This equation considers that the probability Pi at the left side of the equation

should be between zero and one, but the linear vector product xiβ at the right side could include all the real numbers. A

simple method for solving this problem is to use the probability transfer function to remove the distance limits and

model the transferred function as a linear function of the parameters. This conversion occurs in two steps. First, the Pi

probability changes to the chance of success according to Equation 1:

𝑂𝑑𝑑𝑆 =𝑝𝑖

1 − 𝑃𝑖

(1)

In the second step, the logarithm of the above-mentioned equation is taken to obtain the logit or success chance

logarithm (Equation 2):

𝐿𝑜𝑔𝑖𝑡(𝑝𝑖) = 𝐿𝑜𝑔𝑃𝑖

1 − 𝑝𝑖

(2)

The results are quite similar. The reverse transfer function, also called anti-logic, is applied to calculate the

probability in terms of logit (Equation 3):

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932

𝐿𝑜𝑔𝑖𝑡−1(𝑍𝑖) =𝑒𝑧𝑖

1 + 𝑒𝑧𝑖 (2)

The fact that the value of the logit function varies between zero and one, is the first reason for using this function in

the probability problems. The second reason in this regard is the form of this function, so that if we start from negative

infinity and move to the right, by increasing z, the value of f (z) does not change much and remains in the range of

zero until we reach the growth threshold. In this range, the value of the function increases rapidly to approach unity,

and at this time, the increase of z does not have much effect on the increase of the function value. Therefore, the logit

is a transfer function that associates the probabilities in the interval (0, 1) with all the real numbers. The negative logit

represents a less than 50% probability, and the positive logit represents a more than 50% probability. Thus, the logit

model is a general linear model that has a logit transfer function. In other words, the logit of Pi probability, instead of

the probability, follows the linear model [27].

3. Results and Analysis

In this part, accidents data in 2016, 2017, and 2018 were used to identify the variables affecting the accidents

leading to damage, injury, and fatality when encountering the freight vehicles. Then, the data were considered in terms

of statistics and frequency. K–S test, Factor Analysis, Friedman, and Logit test analyses were employed to consider

the variables affecting the severity of accidents.

3.1. Frequency Analysis of the Accidents Results

In this study, there is one dependent variable, i.e. accidents severity, and there are 12 independent variables such as

time of the accident, day of the accident, season of the accident, road function type, road pavement condition, accident

point geometry, lighting status, type of freight vehicle responsible for the accident, age of the faulty driver, type of the

non-faulty vehicle, weather condition, the main cause of the accident, where their frequency analysis is presented in

the following part. Figures 3 to 15 present the frequency of each of the variables used in this study from 2016 to 2018,

split by year.

Figure 3. Accident severity frequency (injury, fatality, and damage) by year

Figure 4. Accidents frequency regarding the time of the accidents by year

0.0

10.0

20.0

30.0

40.0

50.0

60.0

70.0

2016 2017 2018

Freq

uen

cy

Property damage Fatal Injury

0

5

10

15

20

25

30

35

40

45

24--6 6--12 12--18 18--24

Freq

uen

cy

2016

2017

2018

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Figure 5. Accidents frequency regarding the day of the accident by year

Figure 6. Accidents frequency regarding season of the accident by year

Figure 7. Accident frequency regarding the type of highway function year

0

5

10

15

20

25

30

35

40

45

50

2016 2017 2018

Freq

uen

cy

First week

midweek

weekend

0

5

10

15

20

25

30

35

spring summer autumn winter

Freq

uen

cy

2016

2017

2018

0

10

20

30

40

50

60

2016 2017 2018

Freq

uen

cy

rural

arterial

highway

Civil Engineering Journal Vol. 6, No. 5, May, 2020

934

Figure 8. Accidents frequency regarding highway surface condition by year

Figure 9. Accidents frequency regarding accident point geometry by year

Figure 10. Accidents frequency regarding lighting condition by year

0

10

20

30

40

50

60

70

80

90

100

dry wet frost

Freq

uen

cy

2016

2017

2018

0

10

20

30

40

50

60

70

80

90

100

straight curve intersection

Freq

uen

cy

2016

2017

2018

0

10

20

30

40

50

60

70

80

2016 2017 2018

Freq

uen

cy

day

noon

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Figure 11. Accidents frequency regarding the type of the faulty freight vehicle by year

Figure 12. Accidents frequency regarding the type of non-faulty vehicle by year

Figure 13. Accidents frequency regarding weather condition by year

0

10

20

30

40

50

60

70

Pickup truck small truck truck Trailer Carry fuel

Freq

uen

cy

2016

2017

2018

0

10

20

30

40

50

60

70

sedan bus & mini-

bus

heavy

vehicle

motorcycle pedestrians Dealing

with object

Reversal Deviation

from road

Freq

uen

cy

2016

2017

2018

0

10

20

30

40

50

60

70

80

sunny rainy snowy foggy

Freq

uen

cy

2016

2017

2018

Civil Engineering Journal Vol. 6, No. 5, May, 2020

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Figure 14. Accidents frequency regarding the type of the main cause by year

The results of Figures 3 to 14 indicate that more than 50% and 40% of road freight accidents are related to the

damage and injury accidents, respectively. Meanwhile, the accidents resulting in fatality are only about 10% of the

total accidents. Furthermore, the majority of accidents occur between 6 a.m. to 12 a.m and 12 p.m to 6 p.m on

weekdays in autumn. In more than 70% of the cases, the accidents occur on daytime light and sunny weather where

roads are on a tangent with no curve and good dry pavement condition, which prevent the diver from acceleration.

Among all the faulty freight vehicles in the road accidents, pickup trucks and trucks are responsible for 60% and 30%

of the accidents, respectively. The other freight vehicles cause 10% of the road freight accidents.

3.2. Results of Kolmogorov Smirnov Test

Initially, to select an appropriate test to evaluate the data, it is essential to ensure that there is a normal distribution

of statistical data. Thus, K–S test was used to examine whether the distribution is normal. Table 1 summarizes the

results of this test. The results indicate that the test is significant for all three years, namely 2016, 2017, and 2018. As a

result, these variables do not have a normal distribution and nonparametric tests should be used to make deductions.

Table 1. K–S test results for 2016, 2016, and 2018

Significance Statistical Test

Variables 2016 2017 2018

0.00 0.216 0.215 0.214 Time of the accidents

0.00 0.201 0.211 0.201 Season of the accident

0.00 0.210 0.222 0.234 Day of the accident

0.00 0.329 0.278 0.444 Type of highway function

0.00 0.507 0.503 0.267 Highway surface condition

0.00 0.513 0.523 0.331 Accident point geometry

0.00 0.461 0.452 0.394 Lighting condition

0.00 0.399 0.367 0.439 Faulty freight vehicle

0.00 0.269 0.259 0.289 Faulty driver's age

0.00 0.397 0.339 0.518 Non-faulty vehicle

0.00 0.411 0.424 0.462 Weather condition

0.00 0.284 0.284 0.286 Main cause

0

5

10

15

20

25

30

35

2016 2017 2018

Freq

uen

cy

inattention to Precedence failure to Yield

Deviation from road Failure to cargo handling safety

technical problem turning ban

Violation of speed Failure to observe distance

Inability to control

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3.3. Results of Friedman Test

Friedman Test can be utilized to test the equality of variable levels rank. In this study, there are 12 independent

variables for ranking the accident severity. Friedman Test is used to determine the rank of each variable. Table 2

presents the data for each variable, Chi-Square statistics range, degrees of freedom, and sig in order. As sig is less than

5%, H0 is rejected and ranking equality (priority) hypothesis of these 12 factors is not accepted. Table 3 also displays

descriptive statistics which indicate the mean rank of each variable. The smaller the mean rank, the more important is

the corresponding variable.

Table 2. Results of Friedman parameter

Year Number of data Chi-square df Sig.

2016 1152 4565.01 11 0.00

2017 987 3960.32 11 0.00

2018 1189 4702.97 11 0.00

Table 3. Results of Friedman Test for the accidents

Independent variables 2016 2017 2018

Average Rank Average Rank Average Rank

Time of the accidents 9.59 12 9.31 12 9.24 12

Season of the accident 7.65 9 8.34 9 8.21 9

Day of the accident 7.04 7 6.99 7 6.91 6

Type of highway function 8.91 11 8.49 10 8.66 10

Highway surface condition 3.95 2 3.86 2 4.19 2

Accident point geometry 3.88 1 3.63 1 3.66 1

Lighting condition 4.38 3 4.32 3 4.33 3

Faulty freight vehicle 5.81 5 6.02 5 5.69 5

Faulty driver's age 7.35 8 7.15 8 7.05 8

Non-faulty vehicle 6.15 6 6.83 6 7.02 7

Weather condition 4.76 4 4.47 4 4.36 4

Main cause 8.54 11 8.59 11 8.69 11

According to the obtained rankings from Table 3, it can be concluded that the most important variables affecting

road freight accidents in all the three years under study are the accident point geometry (straight, horizontal curve, and

intersections), road pavement status, and lighting status, respectively.

3.4. Exploratory Factor Analysis

It is inevitable to come across with a large number of variables in any study. In order to obtain more accurate

analysis data as well as accomplish more scientific and at the same time practical results, researchers have always been

attempting to reduce the number of variables and establish a new structure for them. Therefore, Factor Analysis is

typically used to achieve this goal. Factor Analysis tries to identify basic variables or factors to explain the correlation

pattern among the observed variables. Factor Analysis plays a pivotal role in identifying the hidden variables or

factors utilizing the observed variables.

When performing Factor Analysis, first one should be sure whether the available data could be used for the analysis

purpose. In other words, it should be determined whether the number of the intended data (sample size and the

relationship between variables) is appropriate for Factor Analysis or not. Consequently, in this study the KMO index

and Bartlett test were utilized to test the referred hypothesis.

Table 4 displays the results for the KMO index and Bartlett test in the present study. The more the index

approximates 1, the more appropriate the intended data will be for Factor Analysis. Similarly, if KMO index is smaller

than 0.5, the Factor Analysis results are not appropriate for the intended data and this analysis should not be used to

interpret the results. Besides, the sig value obtained from the Bartlett Test is smaller than 5% for all the cases and

therefore the assumption that the correlation Matrix is known is rejected.

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Table 4. KMO and Bartlett's test result

Year 2016 2017 2018

Kaiser-Meyer-Olkin Measure of Sampling

Adequacy 0.562 0.516 0.524

Bartlett's Test of

Sphericity

Approx Chi-Square 1201.311 877.590 1187.728

df 66 66 66

Sig 0.000 0.000 0.000

The tables obtained from Factor Analysis consist of two parts. The first part is for special values which determine

the factors that are included in the analysis. Those factors whose specific values are smaller than 1 are excluded from

the analysis. In this study, the factors 1, 2, 3, 4, and 5 which their special values are larger than 1 are included in the

analysis.

The second part indicates the specific values of the factors extracted through the rotation. It should be noticed that

in the rotation of all the remaining factors, a proportion of the total changes, which are explained via these 5 factors, is

taken as fixed (approximately 60%). Tables 5 to 7 present the specific values of road freight accidents from 2016 to

2018, respectively.

Table 5. Specific values of vehicle accidents in 2016

Total Variance Explained

Component Initial Eigenvalues Extraction Sums of Squared Loadings

Total % of Variance Cumulative % Total % of Variance Cumulative %

1 1.892 15.765 15.765 1.892 15.765 15.765

2 1.570 13.084 28.848 1.570 13.084 28.848

3 1.259 10.489 39.338 1.259 10.489 39.338

4 1.168 9.737 49.075 1.168 9.737 49.075

5 1.088 9.064 58.139 1.088 9.064 58.139

6 0.982 8.180 66.319

7 0.903 7.521 73.840

8 0.864 7.197 81.038

9 0.715 5.962 87.000

10 0.639 5.326 92.325

11 0.493 4.105 96.430

12 0.428 3.570 100.000

Extraction Method: Principal Component Analysis.

Table 6. Specific values of vehicle accidents in 2017

Total Variance Explained

Component Initial Eigenvalues Extraction Sums of Squared Loadings

Total % of Variance Cumulative % Total % of Variance Cumulative %

1 1.696 14.134 14.134 1.696 14.134 14.134

2 1.570 13.082 27.216 1.570 13.082 27.216

3 1.266 10.547 37.763 1.266 10.547 37.763

4 1.188 9.896 47.660 1.188 9.896 47.660

5 1.102 9.183 56.842 1.102 9.183 56.842

6 0.994 8.285 65.127

7 0.924 7.699 72.826

8 0.891 7.427 80.254

9 0.743 6.191 86.445

10 0.646 5.384 91.828

11 0.557 4.641 96.470

12 0.424 3.530 100.000

Extraction Method: Principal Component Analysis.

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Table 7. Specific values of vehicle accidents in 2018

Total Variance Explained

Component Initial Eigenvalues Extraction Sums of Squared Loadings

Total % of Variance Cumulative % Total % of Variance Cumulative %

1 1.746 14.551 14.551 1.746 14.551 14.551

2 1.518 12.654 27.205 1.518 12.654 27.205

3 1.378 11.487 38.692 1.378 11.487 38.692

4 1.246 10.383 49.074 1.246 10.383 49.074

5 1.047 8.728 57.802 1.047 8.728 57.802

6 1.000 8.332 66.134

7 0.906 7.553 73.687

8 0.851 7.096 80.782

9 0.722 6.018 86.801

10 0.642 5.353 92.154

11 0.554 4.614 96.768

12 0.388 3.232 100.000

Extraction Method: Principal Component Analysis.

Tables 8 to 10 indicate the rotated matrix of the components from 2016 to 2018, which include the factor loads for

each variable in the factors remaining after rotation. The higher the absolute value of these coefficients in each row,

the more noticeable role the related factor plays in the total changes of the given variable.

Table 8. Component rotation matrix for vehicle accidents in 2016

Component

1 2 3 4 5

Time of the accidents 0.136 0.752 0.364 0.066 0.209

Season of the accident 0.673 -0.136 -0.054 0.231 -0.162

Day of the accident 0.053 -0.054 -0.231 0.016 0.661

Type of highway function 0.055 -0.403 0.573 -0.177 0.223

Highway surface condition 0.747 -0.229 -0.104 -0.071 0.081

Accident point geometry -0.136 -0.076 -0.325 0.418 0.402

Lighting condition 0.390 0.648 0.372 0.100 0.163

Faulty freight vehicle -0.021 -0.374 0.491 0.470 -0.086

Faulty driver's age -0.003 -0.137 0.294 0.269 -0.326

Non faulty vehicle -0.003 0.378 -0.361 0.460 -0.378

Weather condition 0.826 -0.115 -0.158 -0.045 -0.028

Main cause -0.058 -0.168 0.055 0.618 0.283

Table 9. Component rotation matrix for vehicle accidents in 2017

Component

1 2 3 4 5

Time of the accidents -.074 0.699 -0.316 0.285 -0.140

Season of the accident 0.297 -0.058 -0.111 0.154 0.460

Day of the accident -0.146 0.111 -0.069 -0.252 0.344

Type of highway function 0.389 -0.474 -0.171 0.323 -0.397

Highway surface condition 0.824 0.098 0.004 -0.165 0.105

Accident point geometry 0.092 0.124 0.464 0.306 0.331

Lighting condition 0.135 0.654 -0.245 0.404 -0.232

Faulty freight vehicle 0.084 -0.428 0.100 0.642 -0.017

Faulty driver's age -0.155 0.046 0.036 0.433 0.551

Non faulty vehicle -0.085 0.402 0.599 -0.033 -0.092

Weather condition 0.804 0.188 -0.044 -0.166 0.136

Main cause 0.198 0.076 0.688 0.052 -0.298

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940

Table 10. Component rotation matrix for vehicle accidents in 2018

Component

1 2 3 4 5

Time of the accidents 0.014 -0.625 -0.324 0.405 0.039

Season of the accident 0.288 -0.180 0.266 0.004 -0.036

Day of the accident -0.120 -0.025 -0.014 0.064 0.825

Type of highway function 0.088 -0.009 0.749 0.245 0.218

Highway surface condition 0.836 -0.015 0.045 -0.214 -0.007

Accident point geometry 0.183 0.465 -0.283 0.182 0.020

Lighting condition 0.274 -0.602 -0.199 0.434 0.142

Faulty freight vehicle 0.026 0.268 0.460 0.598 -0.058

Faulty driver's age -0.017 -0.022 0.024 0.446 -0.508

Non faulty vehicle 0.175 0.343 -0.539 0.269 0.063

Weather condition 0.844 -0.099 0.056 -0.166 -0.019

Main cause 0.297 0.563 -0.119 0.306 0.174

Extraction Method: Principal Component Analysis.

Based on the Factor Analysis performed for 12 variables affecting the road freight accidents in 2016, as indicated

in Table 8, 5 factors are identified as the major factors. Factor Analysis indicates that variables like the season of

accident, road pavement condition and weather condition are ranked as the first important category of factors;

variables like time of accident and lighting status as the second important category of factors; variables like type of

road and faulty vehicle as the third important category of factors; variables like non-faulty vehicle, accident point

geometry, and the main cause as the fourth important category of factors; and finally weekdays and the age of the

driver are ranked as the fifth important category of factors affecting the severity of road freight accidents in 2016.

Based on the Factor Analysis carried out for 12 variables affecting the road freight accidents in 2017, as indicated

in Table 9, 5 factors are identified as major factors. Factor Analysis shows that the first important category of factors

include variables related to the road pavement status and weather conditions; while, variables like time of the accident

and lighting status are the second most important category of factors; variables like accident point geometry, non-

faulty vehicle, and the main cause are the third most important factor; faulty vehicle as the fourth most important

factor; and the season of accident, weekdays, type of highway, and age of the driver are ranked as the fifth most

important category of factors affecting the severity of road freight accidents in 2017.

Based on the Factor Analysis conducted for 12 variables affecting the road freight accidents in 2018, as presented

in Table 10, five factors are identified as major factors. Factor Analysis indicates that factors like season of the

accident, road pavement status and weather conditions are ranked as the first most important factor; while variables

like time of the accident, lighting status, and main cause as the second most important factor, type of road and non-

faulty vehicle as the third; faulty vehicle as the fourth; and weekdays and the age of driver as the fifth most important

factor affecting the severity of road freight accidents in 2018.

3.5. Road Freight Accidents Severity Model

To establish a model for road freight accident severity, 12 independent variables and 1 dependent variable were

defined. Afterward, they were converted into nominal variables (0 and 1) to be aptly used in SPSS. The dependent

variable, i.e. accident severity, was defined as injury, fatality, and damage accidents. As there were a small number of

fatal accidents, such accidents were categorized as injury accidents. Ultimately, the number of dependent and

independent variables was narrowed down into 2 and 12 respectively. The variables include the type of the faulty

vehicle, type of the road function, the main cause, type of collision, age of the faulty driver, time of the accident, road

pavement status, etc. The enter, backward, and forward methods can be utilized to establish the Logit model. Now, it

should be noticed which of the above-mentioned methods contributes to more appropriate output. In other words, one

should consider which of the above-mentioned methods can present a better model for road freight accidents in Gilan

Province.

To identify such a significant issue, the correct percentage and goodness criteria for the fit model were considered

to identify the fitness of the model. The goodness criterion of the fit model is indicated by R2 parameter. This

parameter shows the percentage of the changes in a given dependent variable determined through Logit independent

variables. Also, the correct percentage criterion determines to what extent the model prediction is correct. In other

words, these two criteria are used to make comparisons between the models and identify a better model. It is

performed in such a way that the more the R2 value approaches unity, a better fit model is established. Similarly, a

higher correct percentage value of a given model, indicates a more powerful model in predicting the accidents.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

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It should be noticed that as in the first method (enter) all the variables simultaneously are input into the equation,

the model lacks enough time to appropriately process the data and select the most significant variables; hence, it

cannot be an appropriate method. Because of this, the forward and backward methods are used to input the data into

Logit equation. Each method which has the highest accuracy in predicting the number of accidents could be adopted as

the best method. In the backward and forward methods, respectively, those variables exit and/or enter where by their

exit or entrance, the minimum change would occur in the value of R2 corresponding to the equation. Likewise, exit or

entrance of the variable leads to improvement or in other words increase of R2 value. This method helps us with

choosing the way of entering the independent variables to be analyzed. Applying various methods provides us with

establishing different equations with the same data and ultimately selecting the best equation.

Table 11 summarizes Logit models in two forward and backward methods. The determiner of the best model is its

degree of accuracy when making predictions. Accordingly, the backward method with its higher degree of correct

percentage in all the cases is selected as the best method to establish Logit model for the severity of road freight

accidents.

Table 11. Summary of the forward and backward methods

2018 2017 2016 Logit method

77.7 74.3 73.2 Forward

78.7 79.9 82.1 Backward

As it was explicated, the present backward method model was selected owing to its high degree of accuracy in

predicting accidents. Hence, this chapter is devoted to introducing the best model. The Chi-Square statistic is used to

determine the effectiveness of dependent variables on independent variables and the fitness of the overall model,

which is comparable with F statistics in ordinary regression analysis. Tables 12 to 14 present the backward method

model coefficients for the vehicle accidents in 2016, 2017, and 2018, respectively. According to these Tables, the Chi-

Square model shows whether the independent variable(s) affects the dependent variable or not. As it is observed, in all

the models the Chi-Square values have zero Sig. Therefore, the independent variables affect the dependent variable

and indicate a high degree of fitness.

Table 12. Backward method model coefficients in 2016

Sig. Df Chi-Square Final Step

0.040 3 8.287 Step

Step 7 0.000 27 630.500 Block

0.000 27 630.500 Model

Table 13. Backward method model coefficients in 2017

Sig. df Chi-Square Final Step

0.240 8 17.65 Step

Step 5 0.000 23 518.975 Block

0.000 23 518.975 Model

Table 14. Backward method model coefficients in 2018

Sig. df Chi-Square Final Step

0.270 2 7.255 Step

Step 7 0.000 30 630.459 Block

0.000 30 630.459 Model

By selecting the backward method and then entering all the selected variables into the process of model

development; and after passing a variety of stages, the ultimate model is obtained through this method. Tables 15 to 17

display the variables entered into the model, using the Wald test, and ultimate Logit model. Wald test considers the

significance of the variables entered into the regression equation and it is comparable with t statistics in normal

regression. Based on these Tables, the Logit model for different years can be presented in this way; the Positive

coefficients of the independent variables indicate their positive relationship with the severity of accidents; while, the

negative coefficients demonstrate their negative relation with the severity of accidents.

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Based on Tables 15 to 17, in 2016, variables like autumn, daylight time, the drivers aging 18 to 60, pickup truck,

and not following safety guidelines in case of freighting goods were the most effective variables on the severity of

road freight accidents. Whereas, in 2017, such variables like rural, trucks and trailers, non-faulty private vehicles,

motors, pedestrians, and no following safety guidelines were the most effective variables. When considering the most

effective variables in 2018, it was determined that the variables of accidents occurring between 12 p.m to 6 p.m, rural,

highways, trailers, non-faulty motorcycles, disregarding front are the main causes, and not following safety guidelines

when freighting are the effective variables on road freight accidents.

Table 15. Logit model variables for vehicle accident severity in 2016

Variables Beta S.E. Wald Sig. Exp (Beta)

Season (Autumn) -0.662 0.255 6.725 0.010 0.516

Daylight (day) -0.462 0.184 6.300 0.012 0.630

faulty driver's age (18-30) -1.300 0.409 10.104 0.001 0.273

faulty driver's age (31-45) -1.371 0.389 12.444 0.000 0.254

faulty driver's age (46-60) -1.289 0.405 10.117 0.001 0.275

faulty freight vehicle (pickup trucks) 1.087 0.510 0.002 0.046 1.024

not following safety guidelines when freighting -1.237 0.629 3.862 0.049 0.290

Constant 3.464 0.000 4.591 0.000 3.195

Table 16. Logit model variables for vehicle accident severity in 2017

Variables Beta S.E. Wald Sig. Exp (Beta)

Rural road 1.500 0.700 4.586 0.032 4.480

Daylight (day) -0.463 0.185 6.243 0.012 0.630

faulty freight vehicle (trucks) 1.757 0.853 4.246 0.039 5.794

faulty freight vehicle (trailers) 0.149 0.379 0.156 0.043 1.161

non-faulty vehicle (private cars) -1.625 0.727 5.001 0.025 0.197

non-faulty vehicle (motorcycles) 3.681 1.023 12.935 0.000 9.687

Pedestrians 3.569 1.239 8.294 0.004 5.464

not following safety guidelines when freighting -1.713 0.633 7.320 0.007 0.180

Constant 1.183 0.879 1.811 0.178 3.265

Table 17. Logit model variables for vehicle accident severity in 2018

Variables Beta S.E. Wald Sig. Exp (Beta)

Rural road -0.701 0.206 11.535 0.001 0.496

Daylight (day) 1.385 0.620 4.990 0.025 3.997

faulty freight vehicle (trucks) 0.800 0.399 4.026 0.045 2.225

faulty freight vehicle (trailers) 1.521 0.916 2.754 0.097 4.575

non-faulty vehicle (private cars) 3.877 0.623 3.781 0.000 4.261

non-faulty vehicle (motorcycles) 0.706 0.329 4.602 0.032 0.454

not following safety guidelines when freighting 0.789- 0.516 2.341 0.026 0.454

Constant -0.379 2.627 0.000 1.000 0.684

The most noticeable points obtained from the type of accidents over the years in this case study reveal the need for

special and more attention of the authorities responsible for road accidents. These include the police and ministry of

road and city planning, rural planners, trailers, motorists, as well as drivers’ following safety guidelines while

freighting loads, all of these variables have been effective on the road freight accidents over the years 2016, 2017, and

2018.

4. Conclusion

Through probing into the data concerning the road freight accidents, the present article focused on analyzing the

effects of various variables on the severity of accidents from 2016 to 2018. Based on the results of the frequency of

Civil Engineering Journal Vol. 6, No. 5, May, 2020

943

variables, it was determined that more than 50% and 40% of road freight accidents were related to the damage and

injury accidents, respectively. The accidents leading to fatality just constituted 10% of the total number of accidents.

The majority of accidents occur between 6 a.m to 12 p.m and 12 p.m to 6 p.m on weekdays in autumn. In more than

70% of the cases, road pavement is dry, the weather is sunny, the route is straight, and it is daytime. Among the faulty

freight vehicles in road accidents, the pickup trucks and trucks are engaged in 60% and 30% of the accidents,

respectively. Other freight vehicles account for 10% of road the freight accidents. Furthermore, according to the

obtained ranking, variables of accident point geometry (tangent, horizontal curve, and intersections), road pavement

status, and lighting status have the highest influence on the road freight accident severity.

By performing the Factor Analysis method, 5 most effective factors on road accidents in 2016, 2017, and 2018

were identified. Moreover, the output of the results obtained from Logit model indicates that the severity of accidents

increases as the result of trailers and motorcycles presence. Moreover, regarding freighting goods in rural areas, some

special measures should be taken. Unfortunately, the main cause of road freight accidents is not following safety

guidelines by the drivers of freighting vehicles. Hence, some strict measures should be adopted as well as more

effective fining strategies should be applied for such drivers.

5. Conflicts of Interest

The authors declare no conflict of interest.

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Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

945

Free Vibration of Tall Buildings using Energy Method and

Hamilton’s Principle

Peyman Rahgozar a*

a M. E. Rinker, Sr. School of Construction Management, University of Florida, P.O. Box115703, Gainesville, FL 32611, USA.

Received 17 February 2020; Accepted 16 April 2020

Abstract

In a framed-tube tall building, shear wall systems are the most efficient structural systems for increasing the lateral load

resistance. A novel and simple mathematical model is developed herein which calculates the natural frequencies of such

tall buildings. The analyses are based on a continuous model, in which a tall building structure is replaced by an idealized

cantilever beam that embodies all relevant structural characteristics. Governing equations and the corresponding eigen-

problem are derived based on the energy method and Hamilton’s principle. Solutions are obtained for three examples;

using the separation of variables technique implemented in MATLAB. The results are compared to SAP2000 full model

analysis; and they indicate reasonable accuracy. The computed natural frequencies for structures 50, 60 and 70 storey

buildings were over-estimate 7, 11 and 14 percent respectively. The computed errors indicate that the proposed method

has acceptable accuracy; and can be used during the initial stages of designing of tall buildings; it is fast and low cost

computational process.

Keywords: Tall Building; Framed Tube; Shear Wall; Free Vibration; Natural Frequency.

1. Introduction

Tall building developments have been rapidly increasing worldwide. One of the most critical issues in tall

buildings is choosing proper structural form to resist lateral loads. Lateral deformation must be severely controlled,

that inhabitant feels comfort and to prevent damages to second-grade structural elements. Another vital point in tall

buildings’ design is the dynamic analysis of these structures that is very important because of their more flexibility and

consequently increases of vibrational amplitude and the fact that the dynamic characteristic of structures is mainly

governed by their natural frequencies [1-2]. Therefore, dynamic parameters calculation of tall buildings is essential for

primary designing. Dynamic parameters such as vibrational frequencies and mode shapes can be calculated by

numerical methods such as finite element. While these numerical methods are used for final designing, approximate

methods are very effective for primary designing. Approximate methods can help the designer in cases such as initial

design when dimensions of some constructional members are not specified, comparison of achieved results with more

advanced numerical methods, and finally specifying of structural dynamic behaviour which leads to better designing.

One of the most ordinary approximate methods for dynamic parameters calculation of tall buildings is “continuum

method” in which the tall building’s structure is substituted by a continuum beam, adopting Euler–Bernoulli or

Timoshenko beam theory as the design tool [3]. Considering different kinds of parameters in the substituted beam can

help the designer to achieve natural frequencies and mode shapes with more accuracy. For resistant of high-rise

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091519

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

946

buildings subjected to lateral loadings, framed tube, rigid frame, braced frame, shear wall or coupled shear walls can

be considered. The framed tube is an economic and ordinary form for wide ranges of tall buildings. The most primary

type of framed tube includes four frame panel vertical on each other; this system consists of closely spaced perimeter

columns tied at each floor level by deep spandrel beams to form a tubular structure. The framed tube structure can be

considered to be composed of: (1) two web panels parallel to the direction of the lateral load, (2) two flange panels

normal to the direction of the lateral load. Framed tube behaviour is similar to a cantilever beam, and the columns in

two parallel sides of the neutral axis function tensile and pressed [4]. Besides, frames parallel with lateral load under

bending resulted from lateral loads indicate shear behaviour [5]. Tavakoli et al. [6-7] studied the seismic performance

of outrigger-belt truss system subjected to the earthquake and blast load using finite element and component-mode

synthesis.

Several methods have been presented to analyze framed tube structures. Coull and Bose (1975) presented a method

based on elasticity theory [8]. Coull and Ahmad (1978) presented a method for the achievement of position changes of

the circumferential frame [9]. Kwan (1994) by using equal orthotropic planes, energy method and elasticity theory,

presented equations for determining stress in columns and also for achievement of lateral deflection of the framed tube

[4]. As the most studies of tall buildings directed toward analysis, Alavi et al. (2018, a, b) proposed simplified

methods which are suitable for the preliminary design of high-rise structures [10-11]. About free vibration of tall

buildings, different types of research have been done by several researchers, that in most of them the vibration of the

structures is modelled as the vibration of a cantilever beam [12-14]. Many researchers have studied fundamental

frequencies of tall buildings [15-17]. Kaviani et al. (2008) carried out an approximate method for determining the

natural periods of multistory buildings subjected to earthquake [16]. In this article, based on a continuum approach

and Hamilton’s principle, a simple mathematical method for calculation of natural frequencies of the combined system

of the framed tube and shear wall is presented. In particular, Mohammadnejad and Haji Kazemi in several research

investigated the natural frequencies of the framed tube structures in more details, considering the effects of shear lag

phenomena [18-20].

There are compound and various structural systems for increasing efficiency of framed tube buildings. A more

uniform distribution of axial stress in flange and web frames, and also a decrease in the values of deflection at the

highest level of structures could be obtained using the mega bracing system [21], shear walls shear core, and also

outrigger-belt trusses in the frame tube structures [22-23]. The system which is considered in this article is a combined

system of the framed tube and shear wall. When framed tube and shear wall system subjected to lateral loads, the

shear wall deforms in bending form with downward concavity and with maximum gradient. Interaction of forces

causes that shear wall to restrain deflection of frames in bases, and framed tube is like a restrain for the shear wall

above structure. Therefore, deflection of the construction decreases. In the recent decade, studies about analysis of free

vibration of the frame with shear wall have been done. Kuang (2001) based on continuum method and D’Alembert’s

principle achieved governing differential equations of free vibration of structures with the symmetrical shear wall [24].

Wang (2005) presented an equation for computing the natural vibration of buildings with coupled shear walls which is

proved to be the fourth-order Sturm–Liouville differential equation, and a hand method for determining the first two

periods of natural vibration of the buildings. Also, to determine the first natural frequency of these structures, a

relation has been suggested [25]. In continuance of previous studies, Bozdogan and Ozturk represented an

approximate method based on the continuum approach and transfer matrix method for free vibration analysis of multi-

bay coupled shear walls [26]. Kamgar and Rahgozar (2019) used energy method as a robust method to compute the

roof displacement and axial forces of columns in tall buildings reinforced with a framed tube and outrigger system

[27].

Although free vibration analysis of framed tube system and shear-walled frame has been studied extensively over

the past few decades, there have been few research efforts related to determining vibrational characteristics of the

combined system of framed tube and shear-wall system. Therefore, to fill in the gap, in this study, a simple analytical

method for calculating natural frequencies of the combined system of framed tube and shear walls is presented. On the

basis of the continuum approach, framed tube and shear walls are replaced by an equivalent cantilever beam located at

the mass center. It should be noted that the first natural frequency of any structure has an important issue in

determining the linear and nonlinear response of structures subjected to the dynamic loads. On the other hand,

calculating the values of natural frequencies of structures using numerical methods is computationally expensive.

Therefore, the main aim of this paper is related to calculate the natural frequency of tall buildings that consist of

framed tube and shear walls using simple analytical methods. The three-dimensional structure is replaced by an

equivalent beam. For this purpose, Hamilton's principle is used to obtain the governing equation of a combined

system. Then the characteristic equation is obtained by applying the boundary conditions. The characteristic equation

is solved to calculate the natural frequency. Several numerical examples are solved, and the results are compared with

those obtained from SAP2000 and other work. Finally, the results show the ability of the proposed method in

comparison with the other methods.

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947

2. Lagrange’s Equations For Combined System Of Framed Tube And Shear Wall

In this section a simple mathematical method for calculation of natural frequencies of a combined system of framed

tube and shear walls is presented based on the works done by Kwan (1994); Malekinejad and Rahgozar (2010) [4, 28].

Kwan (1994) proposed a model for the analysis of framed tube structures; those following assumptions are considered

for modelling the framed tube system by using equivalent orthotropic plates [4]:

The material of the structures is homogenous, isotropic and obeys Hooke’s law.

Spacing of beams and columns are uniform throughout the building height.

3-The floor slabs of tall buildings are not deformable in their planes and have no motion perpendicular to their

planes.

The structure is assumed symmetric in plan and height and cannot twist.

All beams and columns are uniforms along with the building height.

The kinetic (Equation 1) and potential energies (Equation 2) of the considered dynamic system are written as

follows [29]:

2b

0

1K(t) m( x )[ y( x,t )] dx

2 (1)

2 2

0 0

1 1

2 2

b b

P( t ) EI( x )[ y ( x,t )] dx S( x )[ y ( x,t )] dx (2)

In which y(x, t) is displacement and S(x) is the shear stiffness GA(x). In which G is the shear modulus, and A(x) is the

cross-sectional area.

The function A is in the form of L integral, between two arbitrary times of 2

,1

t t .

2 2

1 1

t t

t t

A Hdt (K P)dt (3)

Hamilton’s principle represents that A has a stationary value expressed as A 0 , where is known as the

variational operator.

Hamilton’s principle can be written in the following form [28]:

2

1

t b

nc 1 2t 0

A= H + y dx+ L dt = 0 y = 0 at t = t ,t 0 x b

(4)

Using properties of operator and integration by parts, the following matters result:

Differential equations of motion known as the Lagrange’s equation,

Boundary displacements,

Boundary forces,

Eigenvalue solution form.

H can be determined as follows:

1 1 12 2 2H my EIy Sy2 2 2

(5)

Using the Lagrange equation, Eq. (4) can be rewritten as follows:

2

1

t b

t 0

H H HA= y + y + y F y dx dt = 0

y y y

(6)

At this step, the integrand in Equation 6 should be transformed into one containing only 𝛿𝑦 . Therefore, this

equation is integrated by part, both respect to space and time. After simplification, one can obtain:

Civil Engineering Journal Vol. 6, No. 5, May, 2020

948

0

2

1

t b 2

2t 0

H H H H HA= ( ( )+ ( ) ( ) P) ydx dx+ y y x 0x bx y y t y y yx

H H H H( ( )) y ( ( )) y dtx 0x by x y y x y

(7)

By replacing Equation 5 into Equation 7, the following equation with a series of boundary conditions is derived:

2

2(Sy ) ( EIy ) (my) 0 0 x b

x tx

(8)

Using the method of separation of variable and let y( x,t ) Y( x ) T( t ) , two equations can be obtained. The

frequencies can be obtained from the x-dependent equation [29]:

0Ymω)Y(EIdx

d)Y(S

dx

d 2

2

2

(9)

The Equation 9 after simplifying will be changed as follows by definition the and parameters (Equations 11

and 12).

4 22 2 2

4 2

d Y d Yω Y 0 0 n 1

dn dn (10)

2 2Sb

EI (11)

2 2mb

EI (12)

Values of EI, S and m can be determined by applying the boundary conditions Equations 13 to 16. These boundary

conditions are related to the displacement value at the bottom of the structure (𝑌(𝑛=0)), the value of rotation at the

bottom of the structure (𝑌′(𝑛=0)), shear force ([𝑌" − 𝛾2𝑌′]|𝑛=1) and bending moment (𝑌"|𝑛=1) values at the top of the

structure.

Y 0(n 0) (13)

Y 0(n 0)

(14)

2n 1Y Y 0

(15)

n 1Y 0 (16)

The Y function is considered as follows to obtain a solution for the governing equation (Equation 10):

C pnY(n) e (17)

Therefore, the solutions of the equation (Equation 10) will be as follows by considering Equation 17:

p2 4

2 2 21,2 1 3,4 2ω p = ± B , p = ±i B

2 4

(18)

In which:

4 22 2

1B ω4 2

(19)

4 22 2

2B ω4 2

(20)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

949

To solve the Equation 10, one can rewrite it as follows:

G

H

J

K2

Y(n)

Y(n)p(n,ω)

Y(n)

Y α Y(n) (n)

(21)

1 1 2 2

1 1 1 1 2 2 2 2

2 2 2 21 1 1 1 2 2 2 2

2 2 2 21 2 1 1 2 1 2 1 2 2 1 2

cosh(B n) sinh(B n) cos(B n) sin(B n)

B sinh(B n) B cosh(B n) B sin(B n) B cos(B n)p(n,ω)

B cosh(B n) B sinh(B n) B cos(B n) B sin(B n)

B B sinh(B n) B B cosh(B n) B B sin(B n) B B cos(B n)

(22)

After substituting boundary conditions, a nontrivial solution for Equation 21 can be obtained by setting the

determinant of coefficients to zero.

1 2

2 2 2 21 1 1 1 2 2 2 2

2 2 2 21 2 1 1 2 1 2 1 2 2 1 2

1 0 1 0

0 B 0 B=0

B coshB B sinhB -B cosB -B sinB

B B sinhB B B coshB B B sinB -B B cosB

(23)

Solving Equation 23 using MATLAB software yields:

5 4 2 3 3 2 3 3 21 2 1 2 1 2 1 2 1 2 1 1 2 2

3 3 2 3 3 2 2 4 51 2 1 1 2 2 1 2 1 2 1 2 1 2

B B coshB cosB B B sinhB sinB B B cosh B B B cos B

B B sinh B B B sin B B B sinhB sinhB B B coshB cosB 0

(24)

By solving this equation, the natural frequencies are calculated. For numerical study, EI, S, m, parameters should be

calculated. To calculate G in S GA( x ) Kwan’s relations are used [4].

b s

h

stG=Δ Δ

+Q Q

(25)

In which:

3 2b b c

m c m c

Δ (h - d ) (s - d )h= +( )

Q 12E I s 12E I (26)

22s b c

m sc m sb

Δ (h - d ) (s - d )h= +( )

Q G A s G A (27)

Where bI and cI are moments of inertia of the beam and column respectively, sbA and scA are effective shear areas of

the beam and column, and finally mG is the shear modulus of the material.

The following step by step procedures shown the methodology of the proposed method

Determining the values of mass per unit length (m) along height of structure;

Determining the flexural (EI) and shear stiffness (S) of the structure;

Calculation the and parameter using Equations 11 and 12;

Solving Equation 24 to find the B1 and B2 parameters and finally computing the first natural frequency of the

structure using Equations 19 and 20.

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950

3. Examples and Comparison of Results with Computer Analysis

To verify the accuracy and efficiency of the proposed approximate method, three numerical high-rises symmetric

reinforced concrete buildings which consist of a framed tube and shear walls are presented for determining the natural

frequencies [30]. Then, a comparison is presented between the results in order to evaluate the simplicity and accuracy

of this method. Characteristics of these structures are listed in Table 1, also plan and actual system of tall building are

shown in Figure 1.

Table 1. Geometrical characteristics of structures in plan and height

Figure 1. Plan and actual system of tall building consist of framed tube and shear walls

The elastic characteristics of materials are listed in Table 2.

Table 2. Elastic characteristics

𝑬 (𝑮𝑷𝒂) 𝑮 (𝑮𝑷𝒂) 𝝆𝒄(𝒌𝒈/𝒎𝟑) 𝝂

20 8 400 0.25

Equivalent properties of the buildings are listed in Table 3 based on Kwan’s method in 1994 [4].

Table 3. Equivalent properties of tall buildings

𝑮 (𝑮𝑷𝒂) 𝒕 (𝒎)

1.37 0.21

Flexural (EI) and shear stiffness (S) of the framed tube system and shear walls are calculated as follows:

Number of stories Story’s height Spans length Dimensions of shear-wall Plan’s dimensions

n h(m) wS (m) fS (m) B(m) h(m) t(m) 2a(m) 2b(m)

50 3 3 3 2.2 3 0.3 36 36

60 3 3 3 2.2 3 0.35 42 42

70 3 3 3 2.2 3 0.35 42 42

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951

9 9 9 2

8 8 8 2

50 storey EI = 2 ×10 ×6418.41+ 2 ×10 × 2799.42 = 18435.66 ×10 kg.m

S = 20.59 ×10 +63.36 ×10 = 83.95 ×10 kg.m

9 9 9 2

8 8 8 2

60 storey EI = 2 ×10 ×10372.57 + 2 ×10 ×5556.72 = 31858.58 ×10 kg.m

S = 24.16 ×10 +73.92 ×10 = 98.08 ×10 kg.m

9 9 9 2

8 8 8 2

70 storey EI = 2 ×10 ×15386.45 + 2 ×10 ×8975.68 = 48724.26 ×10 kg.m

S = 28.52 ×10 +85.32 ×10 = 113.84 ×10 kg.m

Where m is mass per unit height of the buildings, which is derived as follows:

7655040050 storey m = = 510336 kg / m

150

11805696060 storey m = = 655872 kg / m

180

17957664070 storey m = = 855127 kg / m

210

By substituting the values of EI, S and m for each of the structure into the Equations 11 and 12, and can be calculated. By substituting their values into Equations 19 and 20, 1B and 2B can be found. Finally, by using Equation

24, natural frequencies are calculated based on a computer program which has been developed in MATLAB for three

numerical examples 50, 60 and 70 storey tall buildings. Comparison of computer analysis results (SAP2000) with the

proposed method are listed in Table 4.

Table 4. Comparison of natural frequencies between SAP2000 and proposed approximate method

Number of stories 𝜸 𝝀 𝝎(𝒓𝒂𝒅/𝒔) Percent of Error in 𝝎

Proposed method SAP2000

50 4.25 3.74 1.93 1.80 7

60 3.15 4.65 1.53 1.37 11

70 2.85 6.12 1.27 1.09 14

The calculated natural frequencies for structures 50, 60 and 70 storey tall buildings have over estimate 7, 11 and 14

percent differences with results of computer analysis (SAP2000). The main source of errors between the proposed

approximate method and SAP2000 may be lead from followings: all closely spaced perimeter columns tied at each

floor level by deep spandrel beams are modelled as a tubular structure, the equivalent elastic properties for GA and EI

and neglecting the effect of shear lag in the approximate method have been used.

Also the results for 60 and 70 storey tall buildings with shear walls are compared with the research carried out by

Rahgozar et al. using B-spline functions [30]. As shown in Table (5), the natural frequencies calculated by the

proposed approximate method are overestimated by 15 and 9 percent for 60 and 70 storey building respectively.

Table 5. Comparison of natural frequencies between proposed approximate method and Rahgozar et al. [30]

Number of stories 𝝎(𝒓𝒂𝒅/𝒔) Percent of Error in 𝝎

Proposed method Rahgozar et al. [30]

60 1.53 1.30 15

70 1.27 1.15 9

4. Conclusion

Natural frequencies and mode-shapes play an important role in structural design of tall buildings. Especially the

first natural mode; since it is the dominant component in response of a tall building to earthquake or wind loading. In

this article, an approximate method for free vibration analysis of the combined system of framed tube and shear walls

was presented. In the proposed method, the structure is modelled as a cantilever hollow box with equivalent structural

characteristics. The governing differential equation was derived by energy method and Hamilton’s principle. Applying

appropriate boundary conditions, natural frequencies of the combined system of framed tube and shear walls were

obtained. Comparing to results from comprehensive finite element models; the proposed method overestimate the first

natural frequency by 7% for the 50-storey, 11% for the 60-storey, and 14% for the 70-storey building. Differences are

within acceptable ranges for a quick estimate. Hence, the proposed method may reliably be used for free vibration

Civil Engineering Journal Vol. 6, No. 5, May, 2020

952

analysis of framed tube tall buildings reinforced by shear walls. The proposed method is simple, accurate, economical,

reliable, and especially suitable for use during the preliminary design; where a large number of structures with

different features need to be analyzed.

5. Conflicts of Interest

The authors declare no conflict of interest.

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[19] Mohammadnejad, Mehrdad, and Hasan Haji Kazemi, “A New and Simple Analytical Approach to Determining the Natural

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Tall Buildings.” Iranian Journal of Structures Engineering 4, No. 2 (2017): 76-88.

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[24] Kuang, Junshang, and S.C. Ng, “Dynamic Coupling of Asymmetric Shear Wall Structures: An Analytical Solution.”

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[27] Kamgar, Reza, and Peyman Rahgozar. “Reducing Static Roof Displacement and Axial Forces of Columns in Tall Buildings

Based on Obtaining the Best Locations for Multi-Rigid Belt Truss Outrigger Systems.” Asian Journal of Civil Engineering 20,

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Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

954

Smart City and Modelling of Its Unorganized Flows Using Cell

Machines

Truong Thanh Trung a*

a Candidate of Law, Faculty of Traffic Police Training, People's Police Academy, Hanoi, Vietnam.

Received 04 December 2019; Accepted 08 March 2020

Abstract

The evolution of the digital economy requires the appropriate infrastructure for administrative management and support

of “Smart City". All this makes it possible to look at the problems of the city in a new way. SMART-city is an integrated

infrastructure, an environment for improving the comfortable life and work of all citizens. In urban traffic flows, there

are obstacles in place where traffic flow is not organized. In these places, special solutions, management measures and

safety criteria are required. Such flows and situations should be simulated. This problem is solved based on flow-

intensive management criteria using situational scenarios. The efficiency of flow management on busy highways requires

the consideration of critical factors. In the present work, such a task is investigated using cell machines, which showed

efficiency in streaming tasks of gas dynamics. A purpose of work and its result is a decline system's complexity and

dimension by means of linearization and reduction of algorithmic complexity. The field of cells is considered. If there’s

an obstacle in the cell then the direction by which the obstacle affects minimally is selected (Stochastic assessment has

been used). System analysis of SMART-city problems is also carried out in this work. Adaptive IT infrastructure,

security, virtualization, risk and the multi-criteria decision-making in an uncertain environment has been analyzed.

Keywords: Simulation; Safe Movement; Smart City; Infrastructure; Management.

1. Introduction

The modern city is the city not only with the developed city traditional infrastructure (Roads, Offices, Service

Institutions, etc.), but also it is the city with the developed IT infrastructure, life activity and comfort. The city which

creates conditions to evolution of digital economy and digital business, health care, municipal management,

educations – "the smart city (SMART-city)" [1], intelligent city. SMART – an abbreviation target words: Specific,

Measurable, Achievable, Relevant, and Time bound ("time limited"). The philosophy of "SMART-city" is necessary

in the metropolis, and also it will provide flexible, technological controllability of the city, its infrastructure [2].

Criteria of the “smart” city are:

1) Digital Technologies;

2) Intellectual Systems;

3) Subsystems of the Electronic State, Government, City Administration;

4) New Interactions, Communications, Intellectual processes of the city structures and "Clever Citizens”

susceptible and ready to innovations of city life.

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091520

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

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955

What city systems should be connected to SMART-city? – All transport logistics, infrastructure, power, police,

ambulance, fire, banking, educational and medical systems, and employment services. Continuous access in the

Internet is necessary.

This definition is widespread, but not the only thing for studying of SMART-city. There are also options:

"intellectual city", "digital city", etc. All of them are closely connected with category "Internet of Things: (IoT)". IoT

is a new stage of society informatization at which practically all objects have access to networks. Networks G5 are

developed for high-quality growth of web connections. The smart home system should be under construction on 3D-

connections "all (x), everywhere (y), at any time (t)". According to forecasts the number of the devices connected in

2020 will exceed 50 billion.

The novelty of this article is to consider unorganized flows with obstacles in the cellular space of a smart city, to

build and study a model of managing such flows using situational scenarios that take into account critical factors (flow

intensity, risk situations, poor visibility, etc.).

2. Research Methodology

IT components and the principles of management of SMART-city are underlain at design, planning of city

infrastructure [3]. "Smart subsystems" of SMART-city allow authenticated citizens to participate in self-government

of the city (health care, education, safety, etc.). Except executive power, it's possible to include subsystems to them

"Clever citizen", "Smart building", "Smart infrastructure", "Smart distribution (smart accounting)", "Smart document

flow", etc. Smart-city Data Base – DPC, situational and call centers, the developed monitoring and adaptive analytics,

etc.

The purpose of subsystems – increase in management efficiency and evolution, decrease in expenses, providing

digital services [4]. For example, signing up in policlinic, the electronic diary of children, public city transport on the

basis of Wi-Fi, GPS and flexible management of it, optimization and management for energy consumption. Safety of

citizens and city structures provide the intellectual video analysis (CCTV), the systems of personal, public and

corporate security. "Smart" security policy, prevention of threats, development of preventive protective measures is

necessary (especially, in places of accumulation of citizens in the city).

SMART infrastructure of the city leans on IT, "clouds" and a Blockchain, Big Data, Social Mining, Data Mining,

and etc. They do the city "more smart", as well as citizens. For example, the Safe Region system of Ramensky district

of Moscow uses 611 video cameras (mass, transport or social assignment). Evolutionary innovations – "intellectual

flows", IT tracking, Eye-tracking (the computerized tracking of a look), neurosystems, bases and the systems of

knowledge, supports of solutions, the chat bots, "talking heads" answering questions of the class FAQ in a natural

language. There are basic concepts of SMART city management:

1) The centralized system – actually one server, it's cheaper, but it is less safe (in the 2016th "dropped" Twitter,

Facebook, PayPal, Amazon because of the DDoS-attack of the cracked coffee-machines);

2) Sets of the interconnected and interdependent servers;

3) Sets of the connected, but almost independent servers, for example, the center of the DPC distributing data to

other systems.

The smart city – evolutionary category, it integrates information systems of the city which quickly, becomes hi-

tech, self-governed thanks to digital technologies. Electronic procedures, functionality of participation in

administration, discussion of city problems and decision-making are available to citizens. Since 2014 there are new

standards of management quality [5, 6], which are quite demanding. The main advantages of "smart city" are:

4) Transport - mobility, reduced travel time;

5) Health care - efficiency of information and reduction of costs;

6) Energetic - comfort and safety;

7) Educational - openness, mobility, individuality;

8) Medical - simplification of access, improvement of quality and information value, quality control;

9) Financial - simplification and improvement of transaction security, variety of systems;

10) Environmental - environmental management, information content for the case;

11) Production and construction - optimization (rationalization) of production, expenses, etc.

The role of SMART-city infrastructure in the development, design of urban infrastructure has so far only been

identified. System analysis and simulation is required [7]. One of the interesting tasks of modeling a transport task for

SMART-city has been considered. It's necessary to develop a policy of safe movement in unorganized high-intensity

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956

flows of the metropolis. Flow-based management has been used. In Malinetsky and Stepantsov (2004) study [8]

similar task is modeled by gas-dynamic task, using the apparatus of cell machines [9], which allows to "escaping"

nonlinear descriptions. Dynamic simulation models are also important [10, 11]. Improving the efficiency of flow

management is a pressing problem on busy urban ways. The measure of the intensity of transport flows is difference,

and the effect of transport is increasing. Reliability of the system, reduction of risks of delivery delays, accidents, non-

optimal traffic parameters are our problems. Traditional economic-mathematical models cannot take into account the

dynamism, complexity, system of risks and uncertainties. Critical factors [12] are often not taken into account:

1) Non-stationary environment (traffic intensity);

2) Heterogeneity of transport conditions;

3) Complicating transport company interactions;

4) Risks of extreme and emergency situations;

5) Dynamic adaptation of transport characteristics;

6) Necessity of simulation calculations of motion parameters taking into account probabilistic data (wind,

visibility, and etc.) [13].

The task using cell machines will be formalized. The field where the movement of unorganized objects takes place

is a cellular. There are insurmountable obstacles on the cell field. The field is defined by an orthogonal grid. In each

cell it’s possible to find a driving transport or stationary obstacle, as well as movement (if there’s no obstacle) "left",

"right", "down", "up" strictly in one direction. If there’s an obstacle in the cell, affects minimally may be evaluated in

a stochastic way. Let 𝜏 be the "view" distance or depth of analysis of the situation in the flow, as well as the choice of

direction with the minimum number of vehicles or obstacles. The field is identified with a pair of matrices (F, V),

where;

𝐹 =∥ 𝑓𝑖𝑗 ∥, 𝑖, 𝑗 ∈ 𝑍, 𝑓𝑖𝑗 ∈ 0,1,2,3,4, 𝑉 =∥ 𝑉𝑖𝑗 ∥, 𝑉𝑖𝑗 ∈ 0,1, (1)

Where; "0" is the absence of transport (F) or obstacle (V), and "1" is the presence of transport. We consider the

vicinity of the von Neumann automata type: neighboring cells (i, j) are considered cells "left" – (i–1, j), "right" – (i+1,

j), "top" – (i, j+1) and "bottom" – (i, j–1). In Figure 1, arrows mark the direction of movement into said cells from the

current cell with coordinates (i, j). Of course, other rules can be adopted, including adaptively customizable movement

moves.

Figure 1. Directions of motion to adjacent points from coordinates (i, j)

Point (cell) (a, b) we’ll designate A, then 𝐴 – inversion of this point, change of the direction on opposite, for

example:

𝐶 = (𝑖, 𝑗), 𝐶 = (𝑖, 𝑗), 𝑊 = (𝑖 − 1, 𝑗), 𝑊 = (𝑖 + 1, 𝑗). (2)

Similarly;

𝐸 = (𝑖 + 1, 𝑗), 𝐸 = (𝑖 − 1, 𝑗), 𝑆 = (𝑖, 𝑗 − 1), 𝑆 = (𝑖, 𝑗 + 1), 𝑁 = (𝑖, 𝑗 + 1), 𝑁 = (𝑖, 𝑗 − 1). (3)

At the same time, conditions are correct for neighboring cells:

𝑓𝑖−1,𝑗 = 𝑓𝑖𝑗(𝑊), 𝑓𝑖+1,𝑗 = 𝑓𝑖𝑗(𝐸), 𝑓𝑖,𝑗−1 = 𝑓𝑖𝑗(𝑆), 𝑓𝑖,𝑗+1 = 𝑓𝑖𝑗(𝑁), 𝑓𝑖𝑗 = 𝑓𝑖𝑗(𝐶). (4)

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957

Similarly for elements of matrix V:

𝑉𝑖−1,𝑗 = 𝑉𝑖𝑗(𝑊), (5)

𝑉𝑖+1,𝑗 = 𝑉𝑖𝑗(𝐸), (6)

𝑉𝑖,𝑗−1 = 𝑉𝑖𝑗(𝑆), (7)

𝑉𝑖,𝑗+1 = 𝑉𝑖𝑗(𝑁), (8)

𝑉𝑖𝑗 = 𝑉𝑖𝑗(𝐶). (9)

It’s necessary to set rules of neighborhood and movement through cells [14-15]. We’ll define these rules as

follows;

1. It’s prohibited to move to a busy cage. Let 𝑃𝑖𝑗(1)

(𝐴) be the probability of moving to A≡(W, E, S, N, C). Then:

𝑃𝑖𝑗(1)(𝐴) =

(1−𝑉𝑖𝑗(𝐴))(1−𝑉𝑖𝑗(𝐴))

4. (10)

2. Analysis of surrounding cells by transport (driver) in the stream. If the adjacent cell is occupied, a ban on

advancing into that cell (according to p.1). The remaining cells are viewed "deep" at the distance 𝜏: we sum the

number of cells, in this direction in the state 1 at the distance 𝜏 and, if there’s a cell with an obstacle among them in

the direction, it and the following cells are considered occupied (prohibition of advance). Probabilities are given by

formulas:

𝑃𝑖𝑗(2)(𝑁) = (1 −

1

𝜏(∑ 𝑓𝑖𝑗+𝑘

𝑑𝑖𝑗

𝑘=1 + 𝜏 − 𝑑𝑖𝑗))𝑃𝑖𝑗(1)

(𝑁), (11)

𝑃𝑖𝑗(2)(𝑆) = (1 −

1

𝜏(∑ 𝑓𝑖𝑗−𝑘

𝑑𝑖𝑗

𝑘=1 + 𝜏 − 𝑑𝑖𝑗))𝑃𝑖𝑗(1)

(𝑆), (12)

𝑃𝑖𝑗(2)(𝐸) = (1 −

1

𝜏(∑ 𝑓𝑖𝑗+𝑘

𝑑𝑖𝑗

𝑘=1 + 𝜏 − 𝑑𝑖𝑗))𝑃𝑖𝑗(1)

(𝐸), (13)

𝑃𝑖𝑗(2)(𝑊) = (1 −

1

𝜏(∑ 𝑓𝑖𝑗−𝑘

𝑑𝑖𝑗

𝑘=1 + 𝜏 − 𝑑𝑖𝑗))𝑃𝑖𝑗(1)

(𝑊), (14)

Where 𝑑𝑖𝑗 is the distance from the cell in question to the nearest obstructed cell in that direction, 𝑃𝑖𝑗(1)

is the

probability of movement (p.1).

3. Movement in flow. We set the flows, increasing the probability of movement in the desired direction. For

example, for the N direction:

𝑃𝑖𝑗(𝑁) = 𝑃𝑖𝑗(2)(𝑁) + 𝑎 min 1 − 𝑃𝑖𝑗

(2)(𝑁), 𝑃𝑖𝑗(2)(𝑆), 𝑃𝑖𝑗

(2)(𝑊), 𝑃𝑖𝑗(2)(𝐸), (15)

𝑃𝑖𝑗(𝑆) = 𝑃𝑖𝑗(2)(𝑆) −

1

3𝑎 𝑚𝑖𝑛 1 − 𝑃𝑖𝑗

(2)(𝑁), 𝑃𝑖𝑗(2)(𝑆), (16)

𝑃𝑖𝑗(𝐸) = 𝑃𝑖𝑗(2)(𝐸) −

1

3𝑎 𝑚𝑖𝑛 1 − 𝑃𝑖𝑗

(2)(𝑁), 𝑃𝑖𝑗(2)(𝐸), (17)

𝑃𝑖𝑗(𝑊) = 𝑃𝑖𝑗(2)(𝑊) −

1

3𝑎 𝑚𝑖𝑛 1 − 𝑃𝑖𝑗

(2)(𝑁), 𝑃𝑖𝑗(2)(𝑊), (18)

Where 0 ≤ a ≤ 1 – aspiration coefficient to advance in this direction in a stream.

4. Configured the field. The configurations of the cell-automatic field are set recurrently:

𝐹𝑛+1 = 𝜑(𝐹𝑛), (19)

𝜑 = 𝜑2°𝜑1 (20)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

958

3. Results and Discussion

The main result of the work is the procedure of situational modeling using the cell machines space of a smart city

defined above. Let's describe this procedure. In composition we will set;

𝜑1(𝑓𝑖𝑗) = ∑ 𝑔𝑖𝑗 𝛼 (𝛼), (21)

𝑔𝑖𝑗(𝛼) = 𝑓𝑖𝑗

(𝛼), 𝛽𝑖𝑗(𝛼) = 𝛼

0, 𝛽𝑖𝑗(𝛼) ≠ 𝛼, (22)

Where 𝛽𝑖𝑗 ≡ 𝑁, 𝑊, 𝐶, 𝐸, 𝑆 if 𝑓𝑖𝑗 ≡ 1, the distribution law is also set, or:

𝑃(𝛽𝑖𝑗 = 𝛼) = 𝑃𝑖𝑗(𝛼), (23)

𝑃(𝛽𝑖𝑗 = 𝐶) = 1 − ∑ 𝑃𝑖𝑗 𝛼≠𝐶 (𝐶) (24)

(The case of a "slow" driver), or:

𝑃(𝛽𝑖𝑗 = 𝛼) =

0, 𝑃𝑖𝑗(𝛾) = 0, ∀𝛾

𝑃𝑖𝑗(𝛼) (∑ 𝑃𝑖𝑗 𝛾≠𝐶 (𝛾))

−1

, 𝑃𝑖𝑗(𝛾) ≠ 0, ∃𝛾 (25)

(Case of "restless" driver). Here 𝛼 ≠ С.

The function 𝜑1 reflects the movement to free cells (up to 4 "applicants" per free space), 𝜑2 - resolves a problem

of "Overpopulation" of cages, this function is set by an algorithm (∀𝑖,𝑗 ):

1) If 𝜑1(𝑓𝑖𝑗) ≤ 1, then go to p.5;

2) If 𝑉𝑖𝑗(𝛼)

𝜑1(𝑓𝑖𝑗(𝛼)

) ≠ 0, then 𝛼 = 𝑅𝑁𝐷;

3) Change the states of the field cells:

𝜑1(𝑓𝑖𝑗(𝛼)

) = 1, 𝜑1(𝑓𝑖𝑗) = 𝜑1(𝑓𝑖𝑗) − 1; (26)

4) proceed to p.2;

5) to put:

𝜑2 (𝜑1(𝑓𝑖𝑗)) = 𝜑1(𝑓𝑖𝑗). (27)

Each transport moves in the selected direction, bypassing obstacles in the direction that are most free to move. For

example, as a conversion 𝜑, it’s possible to take turns or departures aside in case of danger of collision, etc.

The proposed procedure and its possibilities for formalizing and modeling unorganized flows in a smart city has

been discussed. Smart city projects tend to be implemented by deeply integrated systems. They consist of subsystems

with different functional components [16, 17]. In the future, the demographic situation, environmental and economic

needs are taken into account. How smart a city should be is determined not only by the administration, but also by

citizens. This is influenced by their interests, way of thinking, education, age, income [18]. For active activity,

directions with significant connections are selected, forms of partnership, models of interactions (В2В, А2В, B2G and

others), as well as the city IT-platform of public use and the system of assessment of cities are determined [19, 20].

Legal aspects are also taken into account, for example, in Belyaev (2019) and Pinchuk (2019) studies [21, 22]. The

core of "Smart City":

1) Innovation in the real sector;

2) Secure IT infrastructure (for example, Big Data [23]);

3) Developed urban infrastructure (Transport, Utilities, Energy, etc.);

4) Integrated "transparent (Electronic)" control;

5) "Smart citizens";

6) Smart health care, etc.

For SMART-city, everything from health systems to utilities, from energy to energy consumption, from transport

Civil Engineering Journal Vol. 6, No. 5, May, 2020

959

to train stations, from personality safety to social safety, from responding to the darkness of lighting to effective

design of energy consumption of highways, buildings in the city is automated. SMART-city is a near reality. Programs

"Electronic Moscow," Information City, "Electronic Government of Moscow" have been implemented in Moscow, the

program "SMART-city - 2030" of creation of "Smart Capital" by 2030 is being promoted [24]. Moscow is actively

moving to uniformity of platforms, active feedback.

Russia is exploring the potential of "Smart Cities" in 164 medium and large cities (NIU "Higher School of

Economics") [25]. Potential leaders - Moscow, Yekaterinburg, Sochi, Kazan, etc.

The consulting company McKinsey predicts the emergence of more than 600 smart cities in 2020, which will

generate more than 67% of world GDP. Provided that the environment is preserved, energy and economy

management. But it’s necessary to provide forecasts with appropriate models, a base of up-to-date information on

infrastructure. SMART technologies here are a means to achieve a well-equipped urban environment, dialogue the

authorities with the population. "Clever Residents" - informative and creative creators. "Smart city" covers transport

mobility, communal systems, healthcare, education, public safety, finance, trade, production and ecology. The Smart

City Project for all kinds of situations such as those discussed in this work. Unlike other traffic flows modelling (e.g.

[26, 27]), flows are considered unorganized in this study. For such flows, procedures similar to those we have

proposed are effective.

4. Conclusion

The growth of flows and volumes of data of the modern city requires reliability, productivity of its infrastructure

subsystems. For example, in the processing of commercial or housing and communal data, there may be losses in both

directions ("Client-Company"). The advanced IT infrastructure here will benefit, among other things, with the help of

visualization, virtualization and situational modelling. Into account the risks of a multi-criterion decision in conditions

of uncertainty are needed. It’s necessary to organize SMART city's infrastructure and assess the current situation,

adaptive integration of systems and evolution's potential. The flow management in the city requires consideration of

critical factors such as traffic intensity, adaptation of motion parameters, etc.

In the SMART city interaction of subsystems "Citizen", "Building", "Transport", "Logistics", “Housing and

communal services", "Document circulation" and others are based on Big Data, Social Mining, Data Mining and

adaptive analytics. It's necessary for audit the infrastructure of SMART-city to assess the current situation, integration

of administrative solutions and systems, evolution of the city infrastructure. The digital economy is the evolutionary

foundation of SMART-city and a knowledge-based society. By 2025, 5G networks should be deployed in most

Russian cities. Universities will produce 100,000 IT specialists annually. They should support urban SMART-city

infrastructure. The "SMART-city" paradigm is applicable in a metropolis and town. It is flexible, technological,

"smart," effective. It is necessary to get rid of multicriterality, uncertainty (for example, due to unorganized flows of

people and transport) in solving problems of "smart" city. Here, traditional mathematical (e.g., gas-dynamic) models

are complex, requiring complex identification procedures. As in this article, non-classical methods should be used -

cell machines, neural systems, fuzzy logic, situational stochastic modeling, etc.

5. Conflicts of Interest

The authors declare no conflict of interest.

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Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

961

The Effects of Different Shaped Baffle Blocks on the Energy

Dissipation

Nassrin Jassim Hussien Al-Mansori a*

, Thair Jabbar Mizhir Alfatlawi b,

Khalid S. Hashim c, Laith S. Al-Zubaidi

d

a Department of Environment Engineering, University of Babylon, Babylon, Iraq.

b Department of Civil Engineering, University of Babylon, Babylon, Iraq.

c Department of Civil Engineering, Liverpool John Moores University, Liverpool, United Kingdom.

d Ministry of Industry and Minerals, Iraq.

Received 03 December 2019; Accepted 17 March 2020

Abstract

Stilling basins can be defined as energy dissipaters constructed of the irrigation systems. This study aims at investigating

the performance of the new seven baffle blocks design in terms of reducing the dimensions of stilling basins in irrigation

systems. In order to assess the hydraulic efficiency of a new model for baffle block used in stilling basins, a Naval

Research Laboratory (NRL) has conducted. The results of this study demonstrate that the performance of the new baffle

block, in term of hydraulic jump length reduction and hydraulic energy dissipation, it's better than standard blocks.

However, the ratios of the drag resistance attributed to the new baffles block (FB / F2) have been larger than that applied

on the normal block. It was found that the new block dissipates the energy by 9.31% more than the concrete block, and

decreases the length of the hydraulic jump by 38.6% in comparison with the standard blocks. However, the new block

maximizes the drag force ratio by 98.6% in comparison with the standard baffle blocks. The findings indicated that in

terms of energy reduction and dissipation in the length of the hydraulic jump, the new block is superior to the other

kinds.

Keywords: Baffle Blocks; Stilling Basins; Energy Dissipation; Spillway; Hydraulic Jump.

1. Introduction

Stilling basins can be defined as energy dissipaters constructed downstream of the irrigation systems (such as

chutes and spillways). The dimensions of these dissipaters mainly depend on the hydraulic jump characteristics.

Normally, the stilling basins have large dimensions, which means they require large areas and the construction costs

are high. For example, Samadi-Boroujeni et al. (2013) investigated the characteristics of hydraulic jump in a

rectangular cross-section flume over six triangular corrugated beds. Results showed that the folded bed influenced the

conjugate depths of the hop and the water-powered hop length to be reduced by 25% and 54.7% separately [1]. The

creators showed that the amazed beds were superior to the separated ones as far as diminishing the sequent

profundities and length of bounce. As of late, the idea of astound squares has been utilized in cutting edge water

treatment units to blend water and disseminate the unreasonable vitality [2-5].

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091521

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

962

In the skimming flow regime, enhanced energy dissipation has been studied by many researchers who focused on

stepped chutes operating, such as Takahashi and Ohtsu (2012), Hunt et al. (2014) and Chanson (2015) [6-9]. Specific

characteristics of stilling basins Type III at the ends of steeply sloping stepped chutes were investigated by several

researchers, such as Meireles et al.(2014) [10]. The major results of this study are that the hydraulic jump downstream

of a stepped chute stabilized much faster than in a Type I stilling basin. Several empirical methods are available to

provide the sequent depths and Froude number at stilting basins [11-13].

Ellayn and Sun (2012) showed that relative to those with flat floors, the change in the duration of the leap and the

ratio of the longitudinal depths are 30 to 50% and 16.5 to 30% respectively. In this analysis, the importance of

Reynold's number (Re) can be ignored since, as in Abdeen et al. (2015), the viscous force usually has an almost

negligible effect in hydraulic jumping and open path [14, 15]. Ezizah et al. (2012) studied the impact on the hydraulic

jump length of the increase in strength and roughness length parameters [16]. Valero et al. (2015) have recently

numerically analyzed the effects of chute blocks on both the length and stability of the hydraulic jump for both adverse

conditions and design. The model established showed a good ability to replicate the change in the stability and length

of the hydraulic jump [17].

Tiwari et al. (2011) reviewed the impact of end sill on the performance of non-circular pipe outlet models in the

stilling basin. The authors found that under the same flow conditions, the triangular end sill has a better performance

than other end sill types [18]. Some previous studies have also announced comparative perception, Tiwari et al. (2013)

deconstructed the hypothetically and experimentally submerged water-driven bounce framed in an outspread stilling

basin furnished with a sudden drop. The authors have identified that the sill reduces the heaviness of the gate and

operation constrain, and then the gate turns out to be more monetary [19].

Abdelhaleem (2013) performed the prediction of the downstream scour geometry of a Fayoum style weir and the

minimization of the scour using a semi-circular block row. The method discussed in this study is anything but hard to

use as an external part of existing water systems in order to limit downstream scouring of these structures. Including

various statures and positions of puzzle obstructs of various stream situations, a hundred and 53 runs are performed.

An instance of a level floor without confuses was observed for the test program to measure the impact of using the

bewilder docks. The results are analyzed and graphically presented, and the scour parameters were measured with

clear formulae [20].

Mohamed et al. (2015) focused on the experimental study of the impact of the sill over the case study Nagaa

Hammadi regulator stilling basin on the duration of the reverse flow behind the sill, the pace at the end of the stilling

basin, the dissipation of water, the length of the submerged hydraulic hop and the shape of the scour downstream

regulator apron. It is observed that the sill over the stilling basin has a great effect on the characteristic of the flow and

the depth of the local scour shaped downstream operator, especially for the upstream and downstream sill with correct

and slopped faces. It also indicates that, as the submersion ratio and the Froude amount rise, the reverse flow length

downstream sill decreases. Additionally, utilizing sill with correct upstream and slopped downstream face with Ls/

L=0.6 cuts the duration of the submerged hydraulic hop by a total of 59 percent, contributing to a reduction in the

length of the stilling lake. The Regional depth of scouring downstream hydraulic systems has been decreased by 43%

[21]. In addition, Elsaeed et al. (2016) investigated the effect of calming basin shapes with different end-stage heights

on the characteristics of the submerged hydraulic jump and the energy dissipation downstream of the radial gate and

obtained a clear match of the results from the length of the velocity analysis and jump [22].

Abdelhaleem (2017) experimentally investigated the submerged flow through radial gates with and without a gate

sill. He finished up the negative impact of sills under submerged radial gates and legitimized the nearby scour marvels

that occurred promptly downstream the stilling basin of some current submerged radial gates with a gate sill in Egypt

[23].

For submerged hydraulic jumps with cylinders, Alirezaet et al. (2014) studied the time-averaged properties of the

two flow regimes. The mean stream area, the root of the maximum longitudinal velocity and mass vitality dispersal

were investigated and the redirected surface fly (DSJ) method was seen to be more effective than the re-attaching

divider fly (RWJ) stream systems in that the longitudinal portion of the velocity and dispersing the overabundance

vitality of the approaching current; This study obtained a wide range of stream criteria, but the mid-value of quantities

by program was reduced to time-finding. This information could then not be used to evaluate the choppiness spatial

and temporal properties within the two stream structures. Such a knowledge collection; for example, spatial and

transient time arrangement within a submerged hop with squares of the aggressive stream property [24].

An experimental study was conducted by Ibrahim (2017) to explore the impacts of block shapes on the downstream

flow pattern of a radial wall, the tests showed that the blocks had a strong ability to limit the disconnected effect of the

downstream flow pattern [25]. According to the study of Ali et al. (2018), which found that using the baffle block

induced a decrease in the sequential depth ratio, the duration of the hop ratio and the length of the roller, but the

energy dissipation ratio improved [26]. Jaafar et al. (2018) stated that the best performance could be obtained with the

Civil Engineering Journal Vol. 6, No. 5, May, 2020

963

use of two row design of regular USBR baffle blocks with a blockage ratio of 50 and 375 percent, respectively, and at

defined distances by the sequential depth next to the velocity to be spread almost evenly across the basin range [27].

Al-Husseini (2016) demonstrated that the dissipation of flow energy declines as the flow rate rises, and the rough

phase spillway surface is more efficient compared to other low or high flow surfaces [28].

In this context, the current study focused on investigating the performance of the new design of seven baffle blocks

in terms of reducing the dimensions of stilling basins in irrigation systems. Additionally, the performance of these new

baffle blocks has been compared to the performance of standard types of baffle blocks. In order to assess the hydraulic

efficiency of a new design for baffle blocks used in stilling basins, a naval laboratory study was carried out. This

experimental seven blocks performance was contrasted with normal trapezoidal blocks performance. Using the

dimensional analysis technique, the dimensional parameters of the hydraulic performance of the new baffle block and

the drag force used were analyzed.

2. Material and Methods

2.1. The Studied Baffle Blocks

A new seven baffle blocks with standard trapezoidal baffle block have been manufactured, at the University of

Babylon at the Laboratory of the Faculty of Engineering, using local wood according to U.S.B.R recommendations.

Then, these blocks were painted using a waterproof paint to prevent water leakage that could distort their shape. The

studied traditional type of baffle block has a trapezoidal-shaped section with external dimensions of 6.2 cm in width

and 5.5 cm in height. The internal sides are inclined by 9o, which makes the net height of the short side 3.5 cm, Figure

1 (A). The new seven baffle blocks were a V-shaped block with an interior angle of 30o (V30) (vertical and horizontal

angle), a V-shaped block with an interior angle of 20o (V20) (vertical and horizontal angle), a V-shaped block with an

interior angle of 10o (V10) (vertical and horizontal position), and Semi-Cylinder block (SC).

2.2. Experimental Set Up

Experimental work was carried out using a rectangular open tilting flume made from Perspex. The flume is 17.50

m in length, 0.30 m in depth and 0.30 m in width, and has a bed slope of 1:6. It is supplied with a spillway, 0.355 m in

stature, installed at 6.50 m upstream of the flume. Water depth and scour opening have measured by utilizing points

gage, with accurateness of 0.1 mm, mounted on an aluminum frame that could be moved in both directions, vertically

and horizontally along the bed of the flume. The scour gap area and length were accurately measured using a scale

installed on the flume inner wall. The downstream water depth was controlled by a tailgate. The Plan and location of

the baffle of the flume are shown in Figures 1 and 2. It is located after spillway location about 0.5 m.

A) Standard Baffle Block B) V30 C) V20

D) V10 E) SC

Figure 1. The Studied Baffle Blocks

Civil Engineering Journal Vol. 6, No. 5, May, 2020

964

Figure 2. Experimental flume (a) Longitudinal section view and (b) Plan view [31]

2.3. Theoretical Approach

The experimental work was initiated by installing three blocks of each of the seven types of baffle blocks inside the

flume. Then, the following parameters were measured under different discharge rates between (18 and 28) l/s to ensure

an initial Froude number between (6.5 and 9.2):

1) Discharge and Water depth before ogee spillway;

2) Water depth before hydraulic jump (Y1);

3) Water depth after hydraulic jump (Y2);

4) Length of hydraulic jump (Lj);

5) The initial (V1) and final (V2) velocity at Y1 and Y2 respectively.

The following procedure and calculations should be performed:

1. Check the accuracy of both the initial (V1) and final (V2) velocities by applying continuity equation:

𝑉 = 𝑄/𝑌𝐵 (1)

Where: V= initial and final velocity (m/s); Q= discharge (m3/s); B= flume width (m); and Y= depth of flow (m).

2. Compute the velocity heads (hv1, hv2) from the Equation 2:

ℎ𝑣 = 𝑣2/2𝑔 (2)

Where: hv= velocity head; hv1, hv2 before and after hydraulic jump (m), V= velocity (m/s), g= acceleration (m/s2).

3. Compute the initial and final Froude number:

𝐹𝑟1 = 𝑉1/(𝑔𝑌1)0.5 (3)

𝐹𝑟2 = 𝑉2/(𝑔𝑌2)0.5 (4)

Where: Fr1 and Fr2= initial and final Froude number (dimensionless), respectively; V1 and V2= initial and final velocity

(m/s); and Y1 and Y2= Water depth before and after hydraulic jump (m) respectively.

4. Compute the total energy at position Y1 and Y2:

𝐸𝑖 = 𝑌𝑖 + 𝑉𝑖2 /2𝑔 𝑖 = 1, 2 (5)

Where: Ei= kinetic energy at position Y1 and Y2 (m); Yi: Water depth before and after hydraulic jump (m); Vi= initial

and final velocity (m/s).

5. Compute dissipated kinetic energy (ΔE):

𝛥𝐸/𝐸1 = (𝐸1 – 𝐸2)/𝐸1 (6)

Where: ΔE= dissipated kinetic energy (dimensionless); E1= kinetic energy before hydraulic jump (m); and E2= kinetic

energy after hydraulic jump (m)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

965

6. Compute pressure force of water behind hydraulic jump FB:

𝛾.𝑌1

2

2 - 𝛾.

𝑌22

2 -

𝐹𝑏

𝐵 = 𝜌. 𝑉2

2. 𝑌2 - 𝜌. 𝑉12. 𝑌1 (7)

Where: FB: pressure force (kN), γ and ρ are weight and mass density (kN/m3, kg/m

3), respectively.

7. Compute drag force (F2):

𝐹2 = 𝛾 𝑌22/2 (8)

Where: F2= drag force (kN); γ= weight density (kN/m3); Y2 =Water depth after hydraulic jump (m).

Although other variables that influence the drag force it is mainly influenced by the depth of the water. Thus,

during the experimental work, the impact of this factor was considered in the current study. Drag force on baffle

blocks (dissipated kinetic energy (ΔE) is a function of different variables; therefore, a dimensional analysis was

conducted to generate non-dimensional equations, which in turn are functions for the ratio of (ΔE/E1) or (FB/F2)

where flow is turbulent in all laboratory experiments [29, 30]. The following equations show the non-dimensional

variables:

ƒ1(𝐹𝑟1, 𝐹𝑟2, 𝑤/ℎ𝑏 , 𝑋𝑏/ℎ𝑏 , 𝐿𝑗/𝑌1, 𝑡𝑎𝑛 𝜃ℎ) = 0 (9)

ƒ2(𝐹𝑟1, 𝐹𝑟2, 𝑤/ℎ𝑏 , 𝑋𝑏/ℎ𝑏 , 𝐿𝑗/𝑌1, 𝑡𝑎𝑛 𝜃𝑣) = 0 (10)

ƒ3(𝐹𝑟1, 𝐹𝑟2, 𝑤/ℎ𝑏 , 𝑋𝑏/ℎ𝑏 , 𝐿𝑗/𝑌1, 𝑟/ℎ𝑏) = 0 (11)

Where: Fr1= Froude number upstream of the jump; Fr2 = Froude number downstream of the jump, F2 = drag force

(kN), w :width of baffle blocks (m), Xb= distance between baffle blocks and initial of hydraulic jump (m), hb= height

of baffle blocks (m), (Lj)= Length of hydraulic jump (m), Y1= Water depth before hydraulic jump (m), r= radius of

cylinder baffle blocks (m); and (θh, θv)= horizontal and vertical positions, respectively (degrees).

3. Experimental Procedure and Measurement

Seven various forms of baffle piers have been examined downstream of the spillway models as energy dissipators.

There have been a total of 120 races. Consideration was granted to six separate discharges (Q= 19.62, 17.75, 16.33,

12.5, 10.7 and 6.50 l/s). The spillway was used for each release one fixed water depth Y2 downstream. All the models

used were arranged in the flume to have almost a comparable conduit that was roughly 40%. The downstream water

depth is the column area of the perplex wharfs model from the toe of the spillway (X0) and (Y2). The operation courses

and dimensions of the spillway and various flume parts are shown in Figure 1. Point by point and course of action of

the tried and tested amazing models are shown in Figure 2.

Since the hydraulic jump in the stilling basin should be formed steadily for the given discharge, the Froude number

of the approaching stream Fr1 should be confined to prevent the deep waviness and precariousness of the water

surface. Tests of the research facility show that if Fr1 reaches 6.5 to 9.2.

Runs were started with a first feeding backwater until the depth of the downstream water reaches higher than the

desired water depth of some discharge. Then, step by step, the upstream bolstering was started and balanced at that

point. The back end was brought down slowly until the perfect downstream water depth Y2 was collected. When there

were no significant differences in measures of scour gap (it was seen from this research work that a run took around 2

hours to reach a safe condition); the flume delta bolstering valve was shut down. Finally measurements of scour

opening were determined using a point test and scale.

The water level calculations were made in the flume's three dimensional axes. The hydraulic characteristics,

dimensions, and spacing of each individual baffle block are identical for a comparative performance between the

tested baffle blocks. To insure so, downstream water levels are raised, just as the pipeline is held steady to seek. For

each perplex rectangle, the go of the conduit is about 40 percent. The conceptual models of bewildering docks that

have bended surfaces give the impression that they are not constructible, yet they are feasible that they are effectively

created from cement or steel. Again, the aim of this investigation is to propose another bewildering squares, but in

addition to improving each other effectiveness. Figure 3 shows how to calculate the how to calculate hydraulics jump

length. Figure 4 shows experimental procedure of this study

Civil Engineering Journal Vol. 6, No. 5, May, 2020

966

Figure 3. Calculate Hydraulic Jump Length

Figure 4. Selected photos from experimental procedure

4. Results and Discussion

The effect of the form of the baffle blocks on the dissipation of electricity, the ratio of reduction in the duration of

the hydraulic hop, and the ratio of the applied drag force on the baffle blocks was determined using Equations 9 to 11

at Xb/Y1.

4.1. Ratio of Kinetic Energy Dissipation

The experimental work was initiated using the new shapes of blocks (V10, V20, and V30). These V-shaped blocks

were installed in two different positions: horizontally and vertically. For the horizontal position, Figure 5 shows the

relationship between the slope of the V-shaped baffle blocks and the ratio of energy dissipation for Froude numbers

ranging between 6.5 and 9.2. It can be seen from this figure that the larger the slope of the block, the greater the

dissipation of energy for all Froude number values. For example, at Fr= 6.5, the energy dissipation increased from

2.42 to 7.09% as the slope increased from 10 to 30º.

Figure 5. Relationship between slope of the interior angle of the V-shaped baffle blocks and ratio of energy dissipation (in a

horizontal position)

50

60

70

80

90

0 0.17 0.36 0.57

Rat

io o

f en

ergy d

issi

pat

ion

%

Slope of intrior angle of V-shape baffle blocks

Fr=6.5

Fr=7.2

Fr= 7.81

Fr= 8.02

Fr= 8.36

Fr=9.2

Civil Engineering Journal Vol. 6, No. 5, May, 2020

967

A similar relationship between the slope value and the energy dissipation was noticed in the vertical position, as

shown in Figure 6. For instance, at Fr = 6.5, the ratio of energy dissipation increased from 2.62 to 7.93% as the slope

increased from 10 to 30º. However, it can be seen from Figures 5 and 6 that the energy dissipation ratio at the vertical

position was greater than that at the horizontal position. For example, at Fr = 6.5 and slope of 30º, the energy

dissipation ratio increased from 7.09% at the horizontal position to 7.93% at the vertical position. As a compared with

Hayder (2017) which explain the average values of energy dissipation were computed and found to be 68.1% and

60.4% for semi-circular rough elements and stilling basin Type I, respectively [31]. The semi-circular elements model

M1 gave good results of reducing the scour-hole and consequently has good energy dissipation compared to most of

the other models tested by Bestawy (2013) [32].

Figure 6. Relationship between slopes of the interior angle of the V-shaped baffle blocks and ratio of energy dissipation

(in a vertical position)

Again, a similar trend was observed in the semi-cylinder baffle blocks in both vertical and horizontal positions,

where it can be seen from Figure 7 that the energy dissipation ratio increased with the increase in the slope value for

all Fr values. The gap between the vertical and horizontal locations was very minimal because the vertical portion of

the flow will be mirrored at an angle of 180o in the semi-cylinder baffle plates, i.e. the part will fully reflect the flow

path causing high turbulence in the fluid.

According to Abbas et al. (2018) [33] which describes the energy dissipation ratio (ΔE /E1) reduced by using the

adverse slope instead of the horizontal slope, but using the double row of the baffle model (D) at the adverse slope to

increase the efficiency of the stilling basin and to convert the energy dissipation reduction to benefit, the average gain

in the slope (-0.06) ratio (ΔE/E1) reaches about 10.7% when using a double slope model (D) baffle instead of a smooth

horizontal bed. But Jaafar and Maatooq (2018) [34] show that the double rows of the baffle block with a blockage

ratio of 50 % and 37.5 %, respectively, were very successful in improving the hydraulic jump properties and, thus, the

quieting basin efficiency.

Figure 7. The relationship between slopes of the semi-cylinder baffle blocks and ratio of energy dissipation

60

65

70

75

80

85

90

95

0 0.17 0.36 0.57

Rat

io o

f en

ergy d

issi

pat

ion

%

Slope of intrior angle of V-shape baffle blocks

Fr=6.5

Fr=7.2

Fr= 7.81

Fr= 8.02

Fr= 8.36

Fr=8.92

60

65

70

75

80

85

90

95

0 0.17 0.36 0.57

Rat

io o

f en

ergy d

issi

pat

ion

%

Slope of semi-cylinder baffle blocks

Fr=6.5

Fr=7.2

Fr= 7.81

Fr= 8.02

Fr= 8.36

Fr=8.92

Civil Engineering Journal Vol. 6, No. 5, May, 2020

968

4.2. Reduction of Hydraulic Jump Length

Figure 8 shows the relationship between the slopes of the baffle block slopes, in the horizontal position, with the

ratio of Lj/Y1 for Fr ranging between 6.5 and 9.2. A reverse relationship can be noticed between the length of the

hydraulic jump and the baffle block slopes for all the studied values of Fr. For example, at Fr = 6.5 and horizontal

position, the ratio of Lj/Y1 has decreased from about 18 to 9% as the slope increased from 10 to 30º.

Figure 8. The relationship between slopes of the V-shaped baffle blocks, in a horizontal position, and ratio of Lj/Y1

Figure 9 shows the relationship between the slopes of the baffle blocks, in the vertical position, and the ratio of

Lj/Y1 for Fr ranging between 6.5 and 9.2. Again, a reverse relationship can be noticed between the length of the

hydraulic jump and the baffle block slopes for all the studied values of Fr. as a compared with Abbas et al, (2018) the

reduction in the hydraulic jump length is greater than 15% compared to that of the Lozenge type with the same

conditions of Froude number range which was used by Bejestan and Neisi (2009) [33, 35]. As compared with Jaafar

and Maatooq (2018), the average reduction in (Y2/Y1) ratio reaches to 18.3 %, while the average reduction in (Lj/Y1)

ratio reaches to 38.1% when the double baffle model (D) at the adverse slope (-0.06) used instead of the horizontal

smooth bed [34].

Figure 9. The relationship between slopes of the V-shaped baffle blocks, in a vertical position, and ratio of Lj/Y1

0

10

20

30

40

0 0.1 0.2 0.3 0.4 0.5 0.6

Len

gth

rat

io (

Lj/Y

1)

Slope of vertically cut baffle blocks

Fr=6.5 Fr=7.2 Fr= 7.81 Fr= 8.02 Fr= 8.36 Fr=8.92

0

10

20

30

40

0 0.1 0.2 0.3 0.4 0.5 0.6

Len

gth

rat

io (

Lj/Y

1)

Slope of baffle blocks in horizontal position

Fr=6.5 Fr=7.2 Fr= 7.81 Fr= 8.02 Fr= 8.36 Fr=8.92

Civil Engineering Journal Vol. 6, No. 5, May, 2020

969

For the semi-cylinder baffle blocks, Figure 10 shows the relationship between the blocks’ slopes and the ratio of

Lj/Y1 at different Fr values (6.5-9.2). It can be noticed that the relationship is similar to that of the V-shape baffle

blocks for all the studied Fr.

Figure 10. The relationship between slopes of the semi-cylinder baffle blocks and the ratio of Lj/Y1

4.3. The Drag Force Applied on the Baffle Blocks

Figures 11 and 12 shows the relationship between the ratio of FB/F2, at different Fr values, and the slope of the V-

shaped block at horizontal and vertical positions, respectively. These figures indicate a direct relationship between the

ratio of FB/F2 and the slope of the V-shaped baffle blocks at both horizontal and vertical positions. For example, at

horizontal position and Fr= 6.5, the ratio of FB/F2 increased from 43 to 82.6% as the slope increased from 100 to 30

0,

respectively (Figure 11). Similarly, at vertical position and Fr = 6.5, the ratio of FB/F2 increased from 45.7 to 94.6% as

the slope increased from 10 to 30o, respectively (Figure 12).

Figure 11. The relationship between slopes of the V-shaped baffle blocks in a horizontal position and the ratio of FB/F2

0

10

20

30

40

0 0.1 0.2 0.3 0.4 0.5 0.6

Len

gth

rat

io (

Lj/Y

1)

Slope of semi-cylindrical cut baffle blocks

Fr=6.5 Fr=7.2 Fr= 7.81 Fr= 8.02 Fr= 8.36 Fr=8.92

0.01

0.11

0.21

0.31

0.41

0.51

0.61

0.71

0.81

0 0.17 0.36 0.57

Dra

g f

orc

e (F

B/F

2)

Slope of horizontally cut baffle blocks

Fr=6.5 Fr=7.2 Fr= 7.81 Fr= 8.02 Fr= 8.36 Fr=9.2

Civil Engineering Journal Vol. 6, No. 5, May, 2020

970

Figure 12. Slopes of the semi-cylinder baffle blocks, in a vertical position, vs the drag force (FB/F2)

Figure 13 shows the relationship between slopes of the semi-cylinder baffle blocks and the ratio of FB/F2 for Fr

values ranging between 6.5 and 9.2. A direct relationship, for all Fr values, has been noticed between the slope of the

baffle blocks and the ratio of FB/F2.

Figure 13. Slopes of the semi-cylinder baffle blocks, in a horizontal position, vs the drag force (FB/F2)

In addition, it can be noticed that the effect of the flow depth variations along the length of the stilling basins as

shown in Figure 14. Finally, due to the recent development in the sensing technology [36, 37] and application of baffle

plates in different water treatment facilities [38, 39], the authors recommend using sensing technologies to monitor the

behaviour of drag force and water jumps, which will provide very useful information for future studies.

Figure 14. Variation of flow depth along stilling basins

0.01

0.11

0.21

0.31

0.41

0.51

0.61

0.71

0.81

0 0.17 0.36 0.57

Dra

g f

orc

e (F

B/F

2)

Slope of vertically cut baffle blocks

Fr=6.5

Fr=7.2

Fr= 7.81

Fr= 8.02

Fr= 8.36

Fr=8.92

Fr=9.2

0.01

0.11

0.21

0.31

0.41

0.51

0.61

0.71

0.81

0 0.17 0.36 0.57

Dra

g f

orc

e (F

B/F

2)

Slope of semi-cylindrical cut baffle blocks

Fr=6.5

Fr=7.2

Fr= 7.81

Fr= 8.02

Fr= 8.36

Fr=8.92

Fr=9.2

0.06

0.065

0.07

0.075

0.08

0.085

0.09

1 1.1 1.2 1.3 1.4 1.5

Flo

w D

epth

D

Length of Stilling Bains

Civil Engineering Journal Vol. 6, No. 5, May, 2020

971

5. Conclusion

With the change in the cutting angle of the baffle blocks in both horizontal and vertical locations, the amount of

energy dissipation decreases with the frequency of Fr. However, this increase in the vertical position is more than it is

in the horizontal position under the same experimental conditions. Ratio of length of hydraulic jump to the initial

depth (Lj/Y1) is inversely proportional to the cutting angle in both vertical and horizontal positions, and it is directly

proportional to Fr value. Additionally, it has been found that the reduction ratio in hydraulic length in the vertical

position is more than it is in the horizontal position under the same flow conditions. According the rules of energy

conservation, the rate of force increasing should be proportionally to the rate of reductions of jump length. And Ratio

of applied drag force on the baffle blocks (FB/F2) increases with the increase of cutting angle and the initial Fr value.

The hydraulic performance and applied drag force on the semi-cylinder baffle blocks are similar to that of the 30 V-

shaped ones under the same conditions. At the end, in this study drag coefficient values were identified in terms of the

main Froude numbers for various baffle pieces. Results indicated that the drag coefficient values for the vertically cut

blocks in the same flow conditions were smaller than the horizontally cut baffle blocks. However, the average values

of the pressure exerted on the surface of the vertically cut baffle blocks were smaller than on other ones, making them

better than others.

6. Conflicts of Interest

The authors declare no conflict of interest.

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Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

974

Numerical Investigation of Stress Block for High Strength

Concrete Columns

Nizar Assi a*

, Husain Al-Gahtani b, Mohammed A. Al-Osta

b

a Department of Civil Engineering, Birzeit University, Palestine.

b Department of Civil Engineering, King Fahd University of Petroleum & Minerals, Dhahran 31261, Saudi Arabia.

Received 25 December 2019; Accepted 28 March 2020

Abstract

This paper is intended to investigate the stress block for high strength concrete (HSC) using the finite element model

(FEM) and analytical approach. New stress block parameters were proposed for HSC including the stress intensity factor

(α1) and the depth factor (β1) based on basic equilibrium equations. A (3D) finite element modeling was developed for

the columns made of HSC using the comprehensive code ABAQUS. The proposed stress parameters were validated

against the experimental data found in the literature and FEM. Thereafter, the proposed stress block for HSC was used to

generate interaction diagrams of rectangular and circular columns subjected to compression and uniaxial bending. The

effects of the stress block parameters of HSC on the interaction diagrams were demonstrated. The results showed that a

good agreement is obtained between the failure loads using the finite element model and the analytical approach using

the proposed parameters, as well as the achievement of a close agreement with experimental observation. It is concluded

that the use of proposed parameters resulted in a more conservative estimation of the failure load of columns. The effect

of the stress depth factor is considered to be minor compared with the effect of the intensity factor.

Keywords: High Strength Concrete; Stress Block; Column; ABAQUS; Finite Element Model.

1. Introduction

A column is one of the most critical members of a framed structure, the failure of which could lead to a

catastrophic failure of the whole structure. In recent years, high strength concrete (HSC) columns have been widely

used in major construction projects, especially, in high-rise buildings. The advancement in concrete technology and

the development of new types of mineral and chemical admixtures have enabled the production of concrete with

compressive strength exceeding 150 MPa. HSC could lead to smaller member sizes for compression members and

therefore provide considerable savings associated with material costs and a reduction of dead loads. Moreover, due to

the superior durability of HSC [1], a considerable reduction of the maintenance efforts and an increase in the service

life of the structure can be attained as compared to the normal strength concrete (NSC).

The increasing use of HSC has led to concern over the applicability of current design codes and standards.

Although, the latest ACI code provides uniaxial bending interaction curves for up to f′c = 83 MPa, the curves are

based on the stress block for normal concrete. Recent research indicates that the behavior of HSC is different from

NSC in many aspects. The shape of the stress–strain relation of HSC differs from that of NSC [2].

Equivalent rectangular stress blocks have been proposed for HSC by either different standard codes or researchers.

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091522

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

975

The stress block parameters of CSA A23.3 [3], NZS 3101 [4], Ibrahim and MacGregor [5] Bae and Bayrak [6] as

published in ACI SP-293-05, or Ozbakkaloglu and Saatcioglu [7], is recommended for HSC as shown in Table 1.

Table 1. Equivalent rectangular stress block parameters for HSC

Standard / Researcher Stress Block Parameters for HSC, f′c (MPa)

CSA A23.3 [3] α1 = 0.85 − 0.0015 fc

′ ≥ 0.67

β1 = 0.97 − 0.0025 fc′ ≥ 0.67

NZS 3101 [4]

α1 = 0.85 − 0.004(fc′ − 55)

0.85 ≥ α1 ≥ 0.75

β1 = 0.85 − 0.008(fc′ − 30)

0.85 ≥ β1 ≥ 0.65

Ibrahim and MacGregor [5] α1 = 0.85 − 0.00125 fc

′ ≥ 0.725

β1 = 0.95 − 0.0025 fc′ ≥ 0.70

Bae and Bayrak [6]

α1 = 0.85 − 0.004(fc′ − 70)

0.85 ≥ α1 ≥ 0.67

β1 = 0.85 − 0.004(fc′ − 30)

0.85 ≥ β1 ≥ 0.67

Ozbakkaloglu and Saatcioglu [7]

α1 = 0.85 − 0.0014(fc′ − 30)

0.85 ≥ α1 ≥ 0.72

β1 = 0.85 − 0.0020(fc′ − 30)

0.85 ≥ α1 ≥ 0.67

Oztekin et al. [8]

k1 = −0.0012fc′ + 0.805

k3 = −0.002fc′ + 0.964

k1k3 = −0.002fc′ + 0.762

Investigation of stress block parameters for high strength concrete has been performed by some researchers. ACI

code [2], however, suggests stress block parameters for normal concrete only. Many stress-strain relationships for

unconfined high strength concrete are found in the literature. Those relationships were developed based on

experimental work conducted by other researchers. Several stress-strain relationships are found in the literature, but

four of them are selected to adapt stress block for high strength concrete.

Lately, many researchers have suggested a stress block of HSC beams and columns. Al-Kamal [9] proposed a

triangular stress block of HSC beams having a concrete strength over of 55 MPa. The proposed stress block was

validated using 52 tested singly reinforced HSC beams. It was mentioned that the modified rectangular stress block to

find the axial and flexural strength of the HSC column was not likely investigated. Tran et al. [10] developed the

parameters of a stress block for geopolymer concrete (GPC) made of fly-ash and slag which, has a compressive

strength of concrete up to 66 MPa. Rectangular stress-block were proposed based on previously developed curves of

GPC materials and two analytical concrete stress-strain models. Ma et al. [11] carried out theoretical analysis and

experimental tests to study the confined columns under compression and bending moment. In the analytical solution,

an equivalent stress block parameter was developed for a confined HSC section in which the compression stress

distributions were investigated to find the suitable equivalent stress block parameters through using different neutral

axis locations. Hasan et al. [12] experimentally investigated the behavior of HSC and steel fibre HSC column

specimens. The interaction diagrams for specimens were developed based on the equivalent rectangular stress

suggested by CSA A23.3-2014 [13]. Baji and Ronagh [14] carried out statistical study of the concrete rectangular

stress block factors. The Monte Carlo Simulation was used to study the influence of improbability of the stress block

factors on the flexural strength of beams. It is found that any reliability analysis is very sensitive to the change in the

statistical properties of concrete stress block factors. Yang et al. [15] developed a general model of an equivalent stress

block that can be used for both light weight concrete and HSC. Peng et al. [16] proposed equivalent rectangular

concrete stress block factors to design reinforced concrete elements in flexural by incorporating the effects of the

strain gradient. Khadiranaikar and Awati [17] developed an equivalent stress block factor for different concrete

strength through the testing of plain concrete columns, RC eccentrically loaded columns, and beams.

In this paper, new parameters for stress blocks were developed based on equilibrium equations to modify and

extend the current code design curves of HSC columns. Those new parameters will be used to develop interaction

diagrams for columns which were tested experimentally as found in the literature. The interaction diagrams were

validated against experimental data from literature and the results obtained from finite element modelling. In addition,

the effects of the stress block parameters of HSC on the interaction curves were studied. A mathematical code was

developed with the capability of creating design charts similar to those provided by ACI for NSC.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

976

2. Stress-Strain Relationship of HSC

Stress-strain relationships were developed in the literature based on the experimental work done on HSC. One

those models was developed and modified by Hognestad [18]. It is called the modified Hognestad model for HSC. The

modified Hognestad model is given in Equation 1 by:

𝑓𝑐 = 𝑓𝑐′ (k

ϵc

ϵcu− (k − 1) (

ϵc

ϵcu)

2

) (1)

Where; fc: concrete stress corresponding to normal strain ϵc, f′c= standard cylinder strength.

k = 2 −𝑓𝑐

′(MPa)−40

70 for 60 MPa ≤ fc

′ ≤ 94 MPa (2)

ϵcu = [2.2 + 0.015(𝑓𝑐′(MPa) − 40)] × 10−3 for 60 MPa ≤ fc

′ ≤ 94 MPa (3)

Popovics [19] found a stress-strain relationship in Equation 4 for concrete subjected to compression:

𝑓𝑐 =ϵc

ϵ0

𝑛 𝑓𝑐′

(𝑛−1)+(ϵcϵ0

)𝑛 (4)

Where;

ϵ0 = 735(𝑓𝑐′)0.25 ∗ 10−6 (5)

𝑛 = 1.0 + 0.058𝑓𝑐′ (6)

Carreira and Chu [20] proposed another stress strain relationship in Equation 4 for concrete subjected to

compression with the parameters given in Equations 7 to 9:

ϵ0 = (0.71𝑓𝑐′ + 168) ∗ 10−5 (7)

𝑛 =1

1−𝑓𝑐

ϵ0 𝐸𝑖𝑡

(8)

Eit = 0.0736 w1.51(𝑓𝑐′)0.3 (9)

Kumar [21] found another stress-strain relationship in Equation 10 for concrete under compression:

𝑓𝑐 =ϵc

ϵ0(𝑛𝑓𝑐

′) ((𝑛 − 1) + (ϵc

ϵ0)𝑛𝑘)⁄ (10)

Where;

Eci = 3320√fc′ + 6900 (11)

𝑛 = 0.8 +fc′

17 (12)

ϵ0 =𝑛𝑓𝑐

(𝑛−1)𝐸𝑐𝑖 (13)

𝑘 = [k = 1 ϵc ≤ ϵ0

𝑘 = 0.67 +𝑓𝑐

62 ϵc > ϵ0

] (14)

Mertol [22] proposed a stress-strain relationship in Equation 15 for high strength concrete, it is given by:

fc =ϵc

ϵc0

𝑛fc′

(𝑛−1)+(ϵc

ϵc0)𝑛∗𝑘

(15)

𝑛 = 0.310 × 0.145fc + 0.78 (16)

𝑘 = 0.10 × 0.145fc + 1.20 (17)

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977

ϵc0 = 0.0033 − 2 × 0.145 × 10−5 × fc′ (18)

ϵcu = 0.0038 − 4 × 0.145 × 10−5 × fc′ (19)

3. Problem Formulation and Methodology

The objectives of this research work were accomplished following the research approach given by the flowchart

shown in Figure ure 1.

Figure 1. Research methodology flowchart

The actual stress-strain relationships for HSC developed by other researchers such as Hognestad, Popovics, and

Carreira & Chu were utilized to find the best representations for stress block parameters (𝛼1, β1) based on static

equilibrium equations. The generalized stress block and equivalent rectangular stress block are shown in Figure 2.

These parameters are derived based on basic equilibrium Equations 20 to 21 as follows:

Develop a new design aids for HSC columns

Research Objectives & Scope

Literature Review

Stress block for NSC and

HSC

Stress-strain relationships for

NSC and HSC Available design aids for columns subjected

to compression and uniaxial bending for both

NSC and HSC

Experimental works from the

literature Finite Element Modelling

Develop new stress block

parameters for HSC

Validation of Developed Parameters

Generate interaction diagrams for columns subjected to

compression and uniaxial bending for HSC

Conclusions

Civil Engineering Journal Vol. 6, No. 5, May, 2020

978

∫ fc dϵcx0

0= 𝛼1β1fc

′ϵcu (20)

∫ fcϵc dϵcx0

0= 𝛼1β1fc

′ϵcu(1 − 0.5β1)ϵcu (21)

Figure 2. Equivalent rectangular stress block

4. Results and Discussion

4.1. Develop New Stress Block Parameters

In order to obtain the equivalent rectangular stress block shown in Figure , for HSC to be used to evaluate the

strength capacity of the structural element, MATHEMATICA software was utilized to evaluate (𝛼1 and β1) the

parameters for different values of 𝑓𝑐′. The exact values for the stress block parameters (𝛼1 and β1) are obtained from

the equilibrium Equations 20 and 21 and by using the stress-strain models proposed by Hognestad, Popovics and

Carreira and Chu as illustrated in Table 2. The relationships between the obtained values of the α1 and β1 and 𝑓𝑐′

values are shown in Figure 1 to 8. It can be observed that a linear relationship can be used with a correlation

coefficient R2 very close to 1. It can be observed that the Hognestad model shows conservative parameters compared

to the other two models. The other two models overestimate the strength capacity of HSC.

Table 2. Values of 𝜶𝟏𝐚𝐧𝐝 𝛃𝟏for different 𝒇𝒄′ or HSC

𝒇𝒄′ (𝐌𝐏𝐚)

Hognestad Popovics Carreira & Chu

𝛂𝟏 𝛃𝟏 𝛂𝟏 𝛃𝟏 𝛂𝟏 𝛃𝟏

60 0.846 0.734 0.882 0.778 0.991 0.942

70 0.826 0.720 0.880 0.767 0.990 0.938

80 0.807 0.707 0.879 0.756 0.989 0.934

90 0.788 0.694 0.877 0.745 0.988 0.930

100 0.769 0.680 0.875 0.734 0.987 0.926

110 0.750 0.667 0.873 0.723 0.986 0.923

120 0.731 0.653 0.872 0.712 0.985 0.919

Figure 1. 𝜶𝟏 variation vs. concrete strength 𝒇𝒄′ based on the stress-strain model proposed by Carreira and Chu [20]

y = -0.0001x + 0.997

R² = 1

0.984

0.985

0.986

0.987

0.988

0.989

0.99

0.991

0.992

50 70 90 110 130

α1

f'c (MPa)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

979

Figure 2. 𝛃𝟏 variation vs. concrete strength 𝒇𝒄′ based on the stress-strain model proposed by Carreira and Chu [20]

Figure 3. 𝜶𝟏 variation vs. concrete strength 𝒇𝒄′ based on the stress-strain model proposed by Hognestad [18]

Figure 4. 𝛃𝟏 variation vs. concrete strength 𝒇𝒄′ based on the stress-strain model proposed by Hognestad [18]

y = -0.0004x + 0.9647

R² = 0.9987

0.915

0.92

0.925

0.93

0.935

0.94

0.945

50 70 90 110 130

β1

f'c (MPa)

y = -0.0019x + 0.9601

R² = 0.9999

0.72

0.74

0.76

0.78

0.8

0.82

0.84

0.86

50 70 90 110 130

α1

f'c (MPa)

y = -0.0013x + 0.8144

R² = 0.9999

0.64

0.65

0.66

0.67

0.68

0.69

0.7

0.71

0.72

0.73

0.74

50 70 90 110 130

β1

f'c (MPa)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

980

Figure 5. 𝜶𝟏 variation vs. concrete strength 𝒇𝒄′ based on the stress-strain model proposed by Popovics [19]

Figure 6. 𝛃𝟏 variation vs. concrete strength 𝒇𝒄′ based on the stress-strain model proposed by Popovics [19]

4.2. Validation of Developed Parameters

4.2.1. Previous Experimental Works

The three different stress-strain models and their corresponding equivalent stress block parameters are verified with

the experimental data for the reinforced concrete columns found in the literature. Experimental work was conducted

for the four different cases shown in Table 3. Each case represents a certain value of concrete strength ( 𝑓𝑐′). For each

case, a set of experimental data for column capacity (Mn, Pn) is found. The four investigated models are performed in

the developed Mathematica code to create their corresponding interaction curves. The validation of the four models

with the experimental results is shown in Figures 9 to 12.

Case 1: Tested experimentally by Lloyd and Rangan [23];

Case 2: Tested experimentally by Ibrahim and MacGregor [5];

Case 3: Tested experimentally Foster and Attard [24];

Case 4: Tested experimentally by Ibrahim [25].

y = -0.0002x + 0.8923

R² = 0.9931

0.87

0.872

0.874

0.876

0.878

0.88

0.882

0.884

50 70 90 110 130

α1

f'c (MPa)

y = -0.0011x + 0.844

R² = 1

0.7

0.71

0.72

0.73

0.74

0.75

0.76

0.77

0.78

0.79

50 70 90 110 130

β1

f'c (MPa)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

981

The material properties and cross-sectional properties for each case are shown in Table 3. The Experimental data in

Table 5 of the pervious experimental works are used to verify the proposed parameters. Table 4 contains the values of

the stress block parameters corresponding to the four cases and evaluated based on the developed equations of α1 and

β1 in Figures 3 to 8.

Figures 9 to 12 depict the difference between the generated interaction diagrams utilizing the Mathematica code

which, use the stress block parameters in Table 4 and the experimental data. It can be seen that the experimental

results are much closer to the stress-strain model proposed by Hognestad [18] and this model is more conservative.

New parameters for an equivalent stress block can be proposed based on this model; the proposed parameters are

given in Equations 22 and 23 by:

α1 = 0.9601 − 0.0019f’c(MPa) (22)

β1 = 0.8144 − 0.00133f’c(MPa) (23)

Table 3. Properties of the rectangular column sections

Case 1 Case 2 Case 3 Case 4

f’c(MPa) 97 126 90 72

fy(MPa) 400 400 400 400

b (mm) 175 200 175 200

h (mm) 175 300 150 300

0.84 0.60 0.89 0.6

ρ (Steel Percentage) 1.30 1.30% 1.30% 1.30%

: is the ratio of the distance between reinforcement steel on two opposite sides to the cross section length in the corresponding direction.

ρ: is the percentage of steel reinforcements to the concrete in a given column.

Table 4. Corresponding stress block parameters for each case using developed equations in Figures 3 to 8

Case f’c(MPa) Carreira and Chu Popovics Hognestad Bae and Bayrak

Case 1 97 0.987 0.928 0.876 0.737 0.775 0.684 0.742 0.670

Case 2 126 0.984 0.917 0.871 0.705 0.719 0.645 0.670 0.670

Case 3 90 0.988 0.930 0.877 0.745 0.788 0.694 0.770 0.670

Case 4 72 0.990 0.937 0.880 0.765 0.823 0.718 0.842 0.682

Table 5. Experimental results for the tested column corresponding to each case

Case 1 Case 2 Case 3 Case 4

Mn (kN.m) Pn (kN) Mn (kN.m) Pn (kN.m) Mn (kN.m) Pn (kN) Mn (kN.m) Pn (kN)

59.90 742.66 177.72 3746.06 20.53 1606.15 58.33 3207.77

59.11 751.47 180.44 3949.79 20.21 1659.55 108.05 2720.45

60.46 998.04 180.02 4204.89 23.01 1707.75 110.55 2729.55

58.89 965.75 202.89 4406.33 36.50 1353.12 140.34 3015.47

42.79 1969.67 226.41 4403.67 37.25 1374.52 - -

39.69 1928.57 - - 47.12 787.35 - -

- - - - 48.70 822.15 - -

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982

Figure 7. Interaction diagrams corresponding to case 1 using the developed equations of stress block parameters in Figures 3 to 8

Figure 8. Interaction diagrams corresponding to case 2 using the developed equations of stress block parameters in Figures 3 to 8

Figure 9. Interaction diagrams corresponding to case 3 using the developed equations of stress block parameters in Figures 3 to 8

-0.12

0.08

0.28

0.48

0.68

0.88

1.08

0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14

Pn

/fc

Ag

Mn/fc Ag h

Experiment

Hognestad

Carreira & Chu

Popovics

Bae & Bayrak

-0.10

0.10

0.30

0.50

0.70

0.90

1.10

0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14

Pn

/ f

c A

g

Mn/fc Ag h

Experiment

Carreira & Chu

Popovics

Hognestad

Bae & Bayrak

-0.10

0.10

0.30

0.50

0.70

0.90

1.10

0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14 0.16

Pn

/ f

c A

g

Mn / fc Ag h

Experiment

Carreira & Chu

Popovics

Hognestad

Bae & Bayrak

Civil Engineering Journal Vol. 6, No. 5, May, 2020

983

Figure 10. Interaction diagrams corresponding to case 4 using the developed equations of stress block parameters in Figures 3 to 8

4.2.2. Finite Element Model

To verify the proposed developed model, 3-D finite element modeling was developed for the columns with cases

number 3 and 4. The finite element modelling was conducted using the comprehensive code ABAQUS. In this work, a

dynamic explicit method was utilized to avoid the problems of convergence associated by the column concrete under

eccentric loading. Figure 13(a) shows the geometry of the developed model consisting of HSC, longitudinal steel,

transverse steel (ties) and steel plates to apply the load.

Finite Element Types and Interaction

A C3D8R 8-noded linear brick, reduced integration, hourglass control element is used to model both the concrete

and the steel plate at the top and bottom of a column. A 2-noded linear 3D truss (T3D2) element was utilized to model

both the longitudinal and transverse steel rebars. The mesh sizes were investigated in the range of 10 to 30 mm to

obtain the optimum mesh size used from the model. The results showed that based on both the accuracy and a

reasonable running time, the mesh size of 15 mm was used as shown in Figure 13(a).

The interaction between the different parts in the model was represented using different approaches of contact. The

interaction between the concrete and steel rebars was modelled as embedded elements in which the concrete was

considered as the host region. The interaction between the concrete and steel plates on the top and bottom was

represented using a tie interaction (perfect bond).

Boundary Conditions and Load Application

The ends of the columns are modelled in which they are restrained in the transverse directions (Uz = 0). The loads

have been applied as a displacement control at the top and bottom of the steel plates to transfer the load uniformly.

Modelling of HSC and Steel Rebars

The concrete is modelled by using the elastoplastic-damage model developed by Lubliner et al. [26] and extended

by Lee and Fenves [27]. The concrete is modelled by using 8-node solid elements and the steel rebar is modelled by

using truss elements. In this paper, the hardening and softening under compression of the concrete are implemented in

the FE code based on the model developed by Popovics [19]. The model requires this data in the form of stress-

inelastic strain. The plastic damage model parameters implemented in ABAQUS are shown in Table 6. The steel

rebars were modelled as an elastic-plastic material with yield stress given in Table 3.

Table 6. Parameters Used for the Concrete in Plastic Damage Model

Young's Modulus Poisson’s Ratio

Dilation Angle ψ Eccentricity ε fbo/fco K

(MPa) (Degree)

Varied 0.2 36 0 .1 1.16 0.67

-0.10

0.10

0.30

0.50

0.70

0.90

1.10

1.30

0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14 0.16

Pn

/fc

Ag

Mn/fc Ag h

Experiment

Carreira and Chu

Popovics

Hognestad

Bae & Bayrak

Civil Engineering Journal Vol. 6, No. 5, May, 2020

984

Figure 11. Finite element model: (a) Steel rebar; (b) Meshing of column

Verification of the Model

A comparison of the experimental works and the numerical results of the interaction curves was conducted for

cases 3 and 4 and is shown in Figures 14 and 15, respectively. The FE results demonstrated a good agreement with the

experimental ones. It can be seen that the difference between the failure loads obtained from the finite element and the

experimental results of cases 3 and 4 are less than 10%.

Figure 12. Experimental and numerical validation of the interaction diagrams for case 3

Figure 13. Experimental and numerical validation of the interaction diagrams for case 4

-0.10

0.10

0.30

0.50

0.70

0.90

1.10

0.00 0.05 0.10 0.15

Pn

/ f

c' A

g

Mn / fc' Ag h

Experiment

Carreira & Chu

Popovics

Hognestad

Bae & Bayrak

FEM

-0.10

0.10

0.30

0.50

0.70

0.90

1.10

1.30

0.00 0.05 0.10 0.15

Pn

/fc'

Ag

Mn/fc' Ag h

Experiment

Carreira and Chu

Popovics

Hognestad

Bae & Bayrak

FEM

(a) (b)

Civil Engineering Journal Vol. 6, No. 5, May, 2020

985

4.3. Stress Block Parameters Effects on Columns Interaction Charts

Stress block parameters affects the interaction diagrams of columns subjected to uniaxial bending and compression

force. These effects are significant and must be considered. The concrete stress-stain diagram for HSC is not the same

as that for normal concrete. The actual stress block of concrete is idealized to an equivalent rectangular stress block to

simplify analysis and design of structural concrete members. The equivalent rectangular stress block of concrete has

two parameters; the stress intensity parameter (𝛼1) and the stress block depth parameter(𝛽1). The MATHEMATICA

code was developed to generate interaction charts for rectangular and circular columns under the effect of uniaxial

bending and compression force. Using that code to draw interaction diagrams for a typical rectangular reinforced

concrete column with steel on two opposite faces and subjected to uniaxial bending and compression force. The

developed MATHEMATICA code was used to draw the normalized interaction charts for a rectangular column

section with different values of 𝛼1, 𝛽1 and steel percentage (𝜌) as shown in Figures 16 to 18.

Figure 14. Stress block parameters effects on columns interaction charts for 𝝆 = 𝟎. 𝟎𝟏, 𝒇𝒚 = 𝟒𝟏𝟒 MPa, 𝒇𝒄′ = 𝟏𝟏𝟎 MPa

and 𝜸 = 𝟎. 𝟕𝟓

Figure 15. Stress block parameters effects on columns interaction charts for 𝝆 = 𝟎. 𝟎𝟒, 𝒇𝒚 = 𝟒𝟏𝟒 MPa, 𝒇𝒄′ = 𝟏𝟏𝟎 MPa

and 𝜸 = 𝟎. 𝟕𝟓

-0.2

0

0.2

0.4

0.6

0.8

1

0 0.02 0.04 0.06 0.08 0.1 0.12 0.14

Pn

/ f

'c A

g

Mn / f'c Ag h

α1=0.85 & β1=0.85

α1=0.85 & β1=0.65

α1=0.65 & β1=0.85

α1=0.65 & β1=0.65

-0.2

0

0.2

0.4

0.6

0.8

1

1.2

0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 0.16 0.18

Pn

/ f

'c A

g

Mn / f'c Ag h

α1=0.85 & β1=0.85

α1=0.85 & β1=0.65

α1=0.65 & β1=0.85

α1=0.65 & β1=0.65

Civil Engineering Journal Vol. 6, No. 5, May, 2020

986

Figure 16. Stress block parameters effects on columns interaction charts for 𝝆 = 𝟎. 𝟎𝟖, 𝒇𝒚 = 𝟒𝟏𝟒 MPa, 𝒇𝒄′ = 𝟏𝟏𝟎 MPa

and 𝜸 = 𝟎. 𝟕𝟓

The stress intensity factor has major effects in the compression control region as shown in Figures 16 to 18. This

parameter 𝛼1 shifts the interaction diagrams upward and downward when it is equal to 0.85 and 0.65, respectively. The

nominal capacity of the column is higher when the value of 𝛼1 is high. The effect of this factor fades as it moves from

the compression control region to the tension control region. Concrete in the tension control region has very slight

effects and the steel controls the strength in this region. Referring to Figures 16 to 18, the effects of the stress depth

parameter 𝛽1 localized near the balanced point on the interaction diagrams. The effect of 𝛽1 does not appear beyond

the balanced point. Its effect is minor as compared to the effects of 𝛼1.

Finally, these two parameters play an important role in evaluating the nominal capacity of concrete columns

subjected to compression with uniaxial bending. Their values depend mainly on the concrete strength. The American

Concrete Institute (ACI) Code provides relationships to evaluate their values for normal concrete only and the use of

ACI code stress blocks for high strength concrete will overestimate the column capacity. The interaction diagram for

case 3 is developed using an ACI code stress block, a finite element model and the proposed stress block parameters as

shown in Figure 19. It can be seen that a good match for the failure loads is estimated by the FE, ACI code and

proposed stress block parameters.

Figure 17. Interaction diagram for previous case 3 developed using ACI code and proposed stress block parameters

-0.4

-0.2

0

0.2

0.4

0.6

0.8

1

1.2

0 0.05 0.1 0.15 0.2 0.25

Pn

/ f

'c A

g

Mn / f'c Ag h

α1=0.85 & β1=0.85

α1=0.85 & β1=0.65

α1=0.65 & β1=0.85

α1=0.65 & β1=0.65

-0.10

0.10

0.30

0.50

0.70

0.90

1.10

0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14

Pn

/ f

c′ A

g

Mn / fc′ Ag h

Experiment

Porposed Parameters

ACI 2011

FEM

ACI-318-14

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4.4. MATHEMATICA Code

The general MATHEMATICA code has been developed using the MATHEMATICA software. The developed

code is capable of developing interaction diagrams for rectangular or circular columns subjected to compression force

with uniaxial bending. The proposed stress block parameters are used as the default values in the developed code.

Interaction charts for HSC columns subjected to compression with uniaxial bending can be easily found. These

interaction diagrams were not previously found in codes or standards.

Code input:

Concrete strength: 𝑓𝑐′

Yielding strength of steel: 𝑓𝑦

Steel arm to whole depth ratio: γ

Evaluation of column strength

The nominal capacity of the column is evaluated based on the given neutral axis depth. The way the code evaluates

the strength is iterative. Simply assume a neutral axis depth (c) and then evaluate the corresponding column capacity.

Figure 20 shows the stress distribution at a certain value of c.

Figure 18. Typical sketch with some definitions for rectangular cross section

Where: y: Arbitrary distance from the extreme tensile steel; y1: locates the distance to yielding under tension; y2:

locates the distance at which steel starts to yield under compression.

Code output

The output of the code is a graph with eight curves each corresponding to a certain steel percentage starting from

1% and ending with 8%. Figure 21 is a typical code output for the following data: 𝛾ℎ; 𝑓𝑐′ = 80 MPa; 𝑓𝑦 = 400 MPa

and γ = 0.75.

Figure 19. Main output of the developed code corresponding to the above typical example

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5. Conclusions

A new stress block for high strength concrete was developed. Interaction diagrams were generated and validated

with actual experimental data for the reinforced concrete columns found in the literature. The developed interaction

diagrams can be used for the analysis and design of columns. At the end of this research work one can conclude the

following:

A new equivalent rectangular stress block is proposed for HSC. The equivalent rectangular stress proposed/used

by ACI code is for NSC and overestimates the capacity of an HSC column. This overestimation in column

capacity mainly comes from the stress intensity factor of the equivalent rectangular stress block. The stress depth

factor has a minor effect near the balanced point on the column interaction diagram and its effect is considered

minor compared with the intensity factor effect.

The new proposed stress block for high strength concrete has been verified with the experimental data found in

the literature and FEM. Consequently, the calculated nominal capacity of the HSC column reflects the real

capacity of the column and a safe design of columns is obtained.

Among all of the investigated models which represent the stress-strain relations of high strength concrete,

Hognestad’s model is considered to be best one. The stress block parameters, computed using this model, have

given results that are very close to the previous experimental results and FEM, and in addition to being

conservative.

A MATHEMATICA code is proposed to develop interaction charts for HSC columns subjected to compression

with uniaxial bending. The code is used for rectangular and circular sections only.

6. Acknowledgements

The authors would like to thank King Fahd University of Petroleum and Minerals for their technical support.

7. Conflicts of Interest

The authors declare no conflict of interest.

8. References

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and Building Materials 102 (January 2016): 478–485. doi:10.1016/j.conbuildmat.2015.10.194.

[2] ACI-318-14, Building code requirements for structural concrete (ACI 318M-14) : an ACI Standard : Commentary on building

code requirements for structural concrete (ACI 318M-14), American Concrete Institute, (2015). doi:10.14359/11333.

[3] CSA-A23.3, Canadian Standards Association. Design of concrete structures. Mississauga, Ont.: Canadian Standards

Association, 2004.

[4] NZS-3101, Concrete Structures. Part 1: The Design of Concrete Structures, Part 2: Commentary, Wellington, New Zeal. (2006).

[5] Ibrahim, Hisham HH, and James G. MacGregor. "Modification of the ACI rectangular stress block for high-strength concrete."

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[6] Bae, Sungjin, and Oguzhan Bayrak. "Stress block parameters for high-strength concrete members." Structural Journal 100, no. 5

(2003): 626-636.

[7] Ozbakkaloglu, Togay, and Murat Saatcioglu. "Rectangular stress block for high-strength concrete." ACI Structural Journal 101,

no. 4 (2004): 475-483.

[8] Oztekin, Ertekin, Selim Pul, and Metin Husem. “Determination of Rectangular Stress Block Parameters for High Performance

Concrete.” Engineering Structures 25, no. 3 (February 2003): 371–376. doi:10.1016/s0141-0296(02)00172-4.

[9] Al-Kamal, Mustafa Kamal. "Nominal flexural strength of high-strength concrete beams." Advances in concrete construction 7,

no. 1 (2019): 001.

[10] Tran, Tung T., Thong M. Pham, and Hong Hao. “Rectangular Stress-Block Parameters for Fly-Ash and Slag Based

Geopolymer Concrete.” Structures 19 (June 2019): 143–155. doi:10.1016/j.istruc.2019.01.006.

[11] Ma, Chau-Khun, Abdullah Zawawi Awang, and Wahid Omar. “Eccentricity-Based Design Procedure of Confined Columns

under Compression and in-Plane Bending Moment.” Measurement 129 (December 2018): 11–19.

doi:10.1016/j.measurement.2018.07.012.

[12] Hasan, Hayder Alaa, M. Neaz Sheikh, and Muhammad N.S. Hadi. “Performance Evaluation of High Strength Concrete and

Steel Fibre High Strength Concrete Columns Reinforced with GFRP Bars and Helices.” Construction and Building Materials

134 (March 2017): 297–310. doi:10.1016/j.conbuildmat.2016.12.124.

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[13] Canadian Standards Association. "CSA A23. 3-14: Design of Concrete Structures." Canadian Standards Association: Toronto,

ON, Canada (2014).

[14] Baji, Hassan, and Hamid R. Ronagh. "Statistical analysis of the concrete rectangular stress block parameters." In The 2013

World Congress on Advances in Structural Engineering and Mechanics Korea. 2013.

[15] Yang K.-H., Sim J.-I., Kang T.H.-K., “Generalized Equivalent Stress Block Model Considering Varying Concrete

Compressive Strength and Unit Weight.” ACI Structural Journal 110, no. 5 (2013). doi:10.14359/51685832.

[16] Peng, Jun, Johnny Ching Ming Ho, Hoat Joen Pam, and Yuk Lung Wong. “Equivalent Stress Block for Normal-Strength

Concrete Incorporating Strain Gradient Effect.” Magazine of Concrete Research 64, no. 1 (January 2012): 1–19.

doi:10.1680/macr.2012.64.1.1.

[17] Khadiranaikar, R. B., and Mahesh M. Awati. “Concrete Stress Distribution Factors for High-Performance Concrete.” Journal

of Structural Engineering 138, no. 3 (March 2012): 402–415. doi:10.1061/(asce)st.1943-541x.0000465.

[18] Hognestad, Eivind. Study of combined bending and axial load in reinforced concrete members. University of Illinois at Urbana

Champaign, College of Engineering. Engineering Experiment Station., 1951.

[19] Popovics, Sandor. “A Numerical Approach to the Complete Stress-Strain Curve of Concrete.” Cement and Concrete Research

3, no. 5 (September 1973): 583–599. doi:10.1016/0008-8846(73)90096-3.

[20] Carreira, Domingo J., and Kuang-Han Chu. "Stress-strain relationship for plain concrete in compression." In Journal

Proceedings, vol. 82, no. 6, pp. 797-804. 1985.

[21] Kumar, Prabhat. “Effect of Strain Ratio Variation on Equivalent Stress Block Parameters for Normal Weight High Strength

Concrete.” Computers and Concrete 3, no. 1 (February 25, 2006): 17–28. doi:10.12989/cac.2006.3.1.017.

[22] Mertol, Halit Cenan. "Behavior of high-strength concrete members subjected to combined flexure and axial compression

loadings." (2006).

[23] Lloyd, Natalie Anne, and B. Vijaya Rangan. "Studies on high-strength concrete columns under eccentric compression."

Structural Journal 93, no. 6 (1996): 631-638.

[24] S.J. Foster, M.M. Attard, “Experimental Tests on Eccentrically Loaded High Strength Concrete Columns.” ACI Structural

Journal 94, no. 3 (1997). doi:10.14359/481.

[25] Ibrahim, Hisham HH. "Flexural behavior of high strength concrete columns." (1994).

[26] Lubliner, J., J. Oliver, S. Oller, and E. Oñate. “A Plastic-Damage Model for Concrete.” International Journal of Solids and

Structures 25, no. 3 (1989): 299–326. doi:10.1016/0020-7683(89)90050-4.

[27] Lee, Jeeho, and Gregory L. Fenves. "Plastic-damage model for cyclic loading of concrete structures." Journal of engineering

mechanics 124, no. 8 (1998): 892-900. doi:10.1061/(ASCE)0733-9399(1998)124:8(892).

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Appendix I: MATHEMATICA Code

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Available online at www.CivileJournal.org

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Rational Organizational and Technological Solutions for Plastering

Boris Vasilyevich Zhadanovsky a, Vladimir Evgenievich Bazanov

b*

a Candidate of Technical Sciences, Professor, Moscow State University of Civil Engineering, 26 Yaroslavskoye Shosse, Moscow, 129337, Russia.

b Candidate of Technical Sciences, Associate Professor, Moscow State University of Civil Engineering, 26 Yaroslavskoye Shosse,

Moscow, 129337, Russia.

Received 12 December 2019; Accepted 13 April 2020

Abstract

Stucco jobs make a considerable share in the total scope of finishing construction operations. Stucco jobs represent an

intricate technology involving a great number of manual operations. Mechanization of stucco operations allows reducing

labor costs on their performance and increasing labor productivity. This paper is aimed at the selection of optimal

workplace practices during façade stucco jobs using powered tools to treat concrete and brick surfaces of outer walls of

buildings and facilities. The paper discusses organizational and technological solutions in performing façade stucco jobs

including workplace management, workflow process, and equipment and tools utilized. An overview of existent powered

tools for the treatment of concrete and brick surfaces is given; the results of undertaken testing of milling tools for the

treatment of concrete and reinforced concrete structures are analyzed. Based on the study findings, the authors have

concluded that in the improvement of productivity and quality of façade finishing jobs, a great role belongs to correct

(rational) organization of labor using the straightforward segmented workflow, performance of works by specialized

crews of workers, and utilization of high-performance tools and appliances. Different locations and composition of

surfaces being stuccoed require different types of powered tools. The development of new prototypes and the

improvement of existent powered tools for surface treatment allows increasing efficiency and reducing the cost of work.

New options of star inertia mills made of different materials for powered tools equipped with a flexible roll are

suggested.

Keywords: Plastering; Technological Solutions; Mechanized Concrete Processing; Milling Cutters; Inertial Mills; Carbide Tools.

1. Introduction

Finishing work in construction is a laborious and responsible process. They ensure the durability of buildings and

their elements, protecting structures from atmospheric influences, and improve the artistic perception of structures.

Plastering is one of the types of decoration of buildings and structures. Plastering works are widely used both in new

construction and in restoration work to preserve historical and architectural monuments. Plastering is a complex

technology, which uses a large number of manual operations [1, 2]. Despite the development of mechanization, their

volume remains significant and in some cases reaches 60%.

To increase labor productivity, a rational, economically sound organization of labor, the acquisition of brigades and

units by skilled workers, and the use of high-performance equipment, tools, and devices are necessary. These issues

are periodically reviewed and investigated by various authors.

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091523

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

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Different techniques of stucco jobs (both manual and mechanical) have their own advantages, areas of efficient use,

and promising directions for development [3-5]. Stucco job mechanization allows reducing labor costs and improving

performance. Stucco job automation realized through use of stucco robots and the findings of a mathematic model of

automatic stucco job management system are presented in the studies by Bock et al. (2018) [6]. The results of creation

and application of an automated wall stuccoing system are discussed in Shreeranga et al. (2017) [7] and Ankush and

Laukik (2017) [8] studies. The proposed machines include an AC electric motor, gear box, cord mechanism, pulley,

batch tray, and guide paths. The machine model was made with regard to a high-quality brick wall and then checked

by on-line testing.

The quality of façade stucco, the cost and completion dates of jobs depend on the structural and technological

solutions chosen [9], the choice of materials and appropriate equipment for leveling finishing coatings [10, 11].

The influence of various factors on the performance of stuccoworkers teams is investigated using both statistical

and mathematic modeling methods. It was established based on a large number of statistical data [12, 13] that process

efficiency largely depends on the surface type and location, operational procedures and workplace management,

qualifications of team members, and experience of workers. The study utilized the methods of DEA (Data

Envelopment Analysis), regression and correlation analysis, wherein a significant variance in labor productivity at

similar projects was noted. Application of construction process modeling for more precise evaluation of labor

productivity is discussed by Bokor et al. (2019) [14] and Lapidus et al. (2018) [15].

Good quality of outer wall preparation – leveling and buffing with powered tools – is very important in stucco job

performance. For mechanical treatment of reinforced concrete, concrete and stone surfaces, high-performance

machines equipped with diamond and carbide-tipped tools are used [16]. Appearance of new materials and

development of new tools necessitate improving the methods of building structures treatment with abrasive materials

and undertaking studies on structures treatment and principles of calculating parameters of buffing using loose

abrasives [17].

Since January 2020, GOST R 57984-2017 has been implemented in the Russian Federation, identical to the

European standard EN 13914-1: 2005 "Design, preparation and application of external rendering and internal

plastering - Part 1: External rendering". These standards reflect the main technological rules for the preparation and

execution of plastering works. However, organizational issues of work are not considered. Meanwhile, the

organization of work remains one of the important aspects of improving their quality and productivity [18].

This article considers rational organizational and technological solutions for the production of works, including the

organization of the workplace, the sequence of work, the equipment and tools used (traditional and modern). An

overview of existent special equipment and tools for mechanical treatment of concrete and brick surfaces, such as

buffing mills with hard metal stars, has been undertaken. The results of laboratory testing of powered tools for

treatment (leveling) of walls made of different materials are given.

2. Materials and Methods

Production of stucco jobs was organized using the straightforward segmented work flow which is straightforward

performance of works wherein the stucco process is broken down into separate simple operations. The main form of

organizing the work of workers in plastering is the implementation of their specialized units of workers, united in

specialized teams. The production of works is carried out mainly using the flow-split method of labor organization:

continuous production of work with the division of the plastering process into separate simple operations. The

dissected organization of the workflow allows to get high technical and economic indicators (increased labor

productivity and overall acceleration of work due to the specialization and qualification of performers in performing

repeated operations).

The composition of the units and teams of plasterers is formed in each specific case, depending on the nature and

total amount of work on the capture. The required number of workers is determined by the size of the front of work,

which can be provided by the applied means of mechanization, taking into account the level of implementation of

production standards. The qualification of performers must meet the requirements of professional standards for the

quality of individual processes.

Characteristics of star inertia mills for surface treatment were studied in the course of laboratory tests. Test stars

were made with sharp tips or of combined type (with teeth made of hard alloys). Star diameter varied between 35 and

50 mm. The following materials were chosen to fabricate the stars: according to the Russian classification – steel R18

(analogue of EN 1.3355), steel 40Kh (analogue of EN 37Cr4), steel ShKh-15 (analogue of EN 1.3505), as well as

combined reinforced plates from baked hard tungsten-cobalt alloy VK8 (analogue of DIN HG30). The number of mill

rounds changed from 600 rpm to 2200 rpm. Mills were tested on concrete plate samples sized 25×40×4 cm with

compression strength of 10, 15, 20, and 30 MPa. As a coarse concrete filler, lime, marble, and granite were used. In

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the course of tests, dependence of the relief height of treated concrete surface at different mill rpms and star material

wear were studied.

3. Results and Discussion

The construction processes that are part of the plastering of facades (for multi-layer plaster systems) are carried out

in the following technological sequence:

Surface preparation;

Hanging walls and installing beacons;

Reception of the prepared plaster mortar and its transportation to workplaces;

Sequential application of the lower (base) layer of the solution (or several layers);

Application of the upper (decorative-protective) layer of the solution and its grout;

Plastering slopes and ebbs;

Pulling rods with cutting angles;

Care for fresh plaster.

The plastering processes should be distributed between specialized units as follows:

Crew 1 - Preparation of surfaces manually or mechanized;

Crew 2 - The application of the lower layer (s);

Crew 3 - Applying and grouting the upper layer of the solution;

Crew 4 - Plastering slopes and ebbs, pulling rods with cutting angles, cutting rust (if necessary).

The level of qualification of performers in the composition of the crews is determined by the requirements of

professional standards for the implementation of the corresponding labor functions (construction operations).

Depending on the general direction of performing external plastering, the building facades are divided into vertical

(Figure 1a) or horizontal grips (Figure 1b).

(a) (b)

Figure 1. Organization of work on plastering the facade with a breakdown: (a) for vertical captures; (b) for

horizontal captures

When working on vertical grips, the crews are arranged in tiers. The length of the capture on the tier is determined

taking into account the daily production of crews. Simultaneous work on different tiers should be carried out with

strict observance of safety rules for work.

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When the facade is broken down into horizontal grips, the crews are arranged along the entire front of the work

within the grips, and each crew occupies a certain plot. The size of the plots is determined by the daily production of

crews. Work on plots is carried out from the borders of adjacent plots in opposite directions.

In general, work is performed in all cases in the direction from top to bottom (from the cornice to the socle of the

aerial part of the foundation). The procedure for providing work with plaster solutions is determined for each specific

object: a receiving device for processing the finished solution is equipped on an on-site site or a mobile mortar unit

(mortar station) for preparing a solution on-site from a dry commercial plaster mixture or from a mixture of individual

components. For small amounts of work, it is advisable to use small-scale plaster stations (mortar mixers) to prepare a

solution from plaster mixtures or to perform manual mixing with the help of an electric tool. Next, the sequence,

methods and techniques of work in plastering the facade were considered.

Crews 1 performs mechanized cleaning of the surface from dust, dirt, salt deposits, felling of flows, sealing of

sinks, notching of smooth concrete and stone walls, hanging surfaces. Dust from the surface of the facade is removed

by a jet of compressed air from the compressor, flushing is carried out with water.

Cutting down of flows and notching of concrete and stone surfaces on small areas is made manually with the help

of chisels and hammers, scarpels. On large areas, mechanized tools are used: pneumatic and electric grinders, electric

brushes, sandblasting machines, cleaning cutters [19, 20].

The entire surface and elements of the facade are verified to determine the degree of deviation of the surfaces and

faces from the horizontal and vertical. Reconciliation of vertical (side) slopes of window openings located on the same

axis is performed by hanging them with a plumb line lowered through all floors from the upper floor window. The

cord along the lines of their faces along the facade is stretched to reconcile the location of the upper slopes and drains

of window openings (Figure 2).

Figure 2. Sagging facade

The places of mismatch of slopes and drains with the corresponding vertical and horizontal lines are corrected by

cutting the protruding parts of the bases or by increasing the thickness of the plaster within the permissible limits

(GOST R 57984—2017 / EN 13914-1-2005).

If necessary (for example, when plastering the facade for painting), crew 1 also performs the installation of

beacons. Lighthouses are installed at all corners of the facade and on the sides of window openings. With a significant

width of window openings and piers, additional beacons are installed so that the leveling of the solution between them

can be done using working rules with a length of 1.5 - 2.0 m. The thickness of the beacons should correspond to the

thickness of the lower (base) layer of plaster.

Crew 2 applies the base layer of plaster (from one or more layers) sequentially as the underlying layer sets. The

solution is applied mechanically or manually.

The surface of the lower layer is well compacted and leveled. If necessary, the surface of the lower layer is cut

horizontally with wavy grooves with a depth of at least 3 mm. Cement window sills are performed manually using the

appropriate tool, with further installation of metal sinks.

Crew 3 before applying the decorative upper layer produces (if necessary) wetting the lower layer by spraying

water. The top coat is applied mechanically.

Crew 4 performs part of the work before the work of crews 2 and 3. The traction of belts, platbands and pilasters is

carried out before plastering smooth surfaces. The slopes of the openings are plastered before applying the top layer of

plaster. At the same time, the joints of the plaster are made not on the front surface of the wall, but on the edges of the

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slopes. Profiles of cornices, belts, platbands are first stretched in the lower base layer, and after hardening, the upper

layer is applied. Templates are used to perform plaster rods.

In the production of external plastering works, the team is equipped with a set of appropriate mechanisms, tools,

devices and equipment.

In general, the plasterers team uses the following set of equipment, tools, and devices (in parentheses are

references to state standards of the Russian Federation for the respective products):

a) Equipment

Plaster Station - for the preparation, processing, transportation and application of the solution during its delivery

over a considerable distance;

b) Mechanized tools

Electric hammer or pneumatic hammer - for cleaning the surface;

Milling cutter - for leveling the surface,

Furrower - for cutting furrows or grooves,

Trowels - for grouting the top layer;

Mixer - for manual preparation (kneading) of a plaster mortar;

Plaster pneumatic hopper bucket (mortar) - for applying mortar to the wall.

c) Hand tools

Falcon duralumin - for applying the solution to the surface and leveling it; duralumin Falcon – for applying the

solution to the surface and leveling it;

Rule (wood or aluminum profile rail) (GOST R 58519-2019. Rules, graters and graters. Technical conditions /

GOST R 58519-2019. Darbies, floats and semifioats. Specifications) - for leveling the bottom layer;

Large grater 1.2 × 0.11 m - for leveling the solution;

Wooden half-glass;

Small grater 0.35x0.05 m - for cutting corners;

Wooden grater - for leveling the lower and upper layers;

Felt grater - for grouting the top layer;

A steel smoothing iron small and large - for leveling and smoothing the top layer;

Plaster trowel (GOST 9533-81 Trowels, tuck pointing tools. Specifications) - for spraying the mortar;

Mortar shovel (GOST 19596-87. Shovels. Specifications) - for shoveling and supplying a solution;

Plastering construction hammer (GOST 11042-90 Construction steel hammers. Specifications);

Brush (GOST 10597-87 Painting brushes. Specifications) - for spraying the plaster layer with water during

grouting, washing contaminated surfaces, washing tools;

Metalwork chisel 10 × 60º, 20x60º (GOST 7211-86. Cold chisels. Specifications) - for cutting down influxes;

Scarpel - for cutting down influxes;

Double-edged hammer - for cutting sagging;

Cutting type OSH-1 (GOST 9533-81 Trowels, tuck pointing tools. Specifications) - for cutting corners;

Manual steel brushes - for cleaning surfaces;

Plastering bucket - for applying mortar.

d) Measuring instruments

A plumb weighing 400 - 600 g (GOST 7948-80 Steel construction plumb-lines. Specifications) - to check the

verticality of structural elements;

Building level 300 - 700 mm long (GOST 9416-83 Building levels. Specifications) - to check the horizontal

structure;

Standard cone (GOST 5802 - 86 Mortars. Test methods) - to check the mobility of the solution;

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Plumb rail (GOST 5802-86 System of ensuring geometric parameters accuracy in building. Rules for measuring

parameters of buildings and works) - for quality control of plaster;

Wooden square - for cutting corners;

A metal square with a mobile bar - for marking slopes;

Measuring metal tape measure up to 20 m long (GOST 7502-98 measuring metal tapes. Specifications) - for

measurements.

e) Devices

Plaster nozzles - for applying solutions;

Plaster corners - metal products for the formation of external corners;

Building mesh (GOST 3826-82. Wire cloth nets with square mesh. Specifications) - mounted between beacons

when a thick layer of mixture is applied on an uneven surface;

Paint net - the material is laid on the wall with shallow cracks when applying a thin layer of plaster.

f) Production inventory

Plaster box - for storing solutions in the workplace;

Buckets - for storing water;

Rubber hoses and hoses with a length of 15 - 20 m - for transporting solutions.

The use of special equipment makes it easier to carry out work, shorten the time for their completion and ensure a

high-quality result. Before finishing layers, including stucco, are applied onto concrete and reinforced concrete

structures, their surfaces must be stripped of fins and other defects which it is expedient to do with powered tools. An

effective way of leveling concrete and cleaning surfaces is the use of milling cutters (for example, the production of

Metabo, Festool). As the working body, milling heads (Figure 3) are used, equipped with solid metal milling

sprockets. Asterisks can be produced in various types: with a pointed tooth (for removing hard plaster, concrete) and

with a flat tooth (for removing soft plaster, fresh concrete).

(a)

(b)

Figure 3. Milling cutter for leveling concrete: (a) manufactured by Metabo (www.metabo.com); (b) production of

Festool (www.ftool.ru)

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For textured finishing of vertical (facades) and horizontal elements of buildings and structures, in addition to the

considered milling cutters with inertial carbide sprockets, shank cutters or groove cutters can be used (Figure 4). They

allow to make notches or cut grooves with a variable width (3 - 40 mm) and a cutting depth (10 - 40 mm).

Figure 4. Wall chaser (sidex.ru)

Along with the considered equipment for removing sagging and uneven walls, it seems rational to use mechanized

tools with a flexible shaft. As a drive, electric grinders with a flexible shaft can be used (such as IE-8201A machines

previously manufactured in Russia - Figure 5) or an inertial vibrator drive for compacting concrete.

Figure 5. Electric Hand Grinders with Flexible Shaft

As a working element, sprocket-shaped inertial mills are used. Asterisks can be spiky (Figure 6a) and combined

with carbide teeth (Figure 6b).

(a) (b)

Figure 6. Schemes of inertial mills: (a) Spiky; (b) Carbide teeth

The location of the sprockets on the shafts of the centrifugal mill can be ordinary or staggered (Figure 7).

(a) (b)

Figure 7. The arrangement of the sprockets on the shaft of the centrifugal mill: (a) Ordinary location; (b) Chess

To evaluate tool productivity and efficiency in outer wall treatment, laboratory tests of mills made of different

materials were carried out: according to the Russian classification, steel R18 (analogue of EN 1.3355), steel 40Kh

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1004

(analogue of EN 37Cr4), steel ShKh-15 (analogue EN 1.3505), as well as combined reinforced plates from sintered

hard tungsten-cobalt alloy VK8 (analogue to DIN HG30). The minimum milling performance was shown at a

frequency of 600 rpm. With increasing speed, productivity increased (Figure 8).

Figure 8. The dependence of the height of the relief hр the machined concrete surface on the peripheral speed n of the

rotation of the cutter: (1) concrete with marble aggregate 15 MPa; (2) concrete with limestone aggregate 20 MPa; (3)

concrete with granite aggregate 30 MPa.

The best results were obtained with a sprocket diameter of 45 mm and a tooth pitch of 11 mm (13 teeth around the

circumference). It was also found that when the distance between the acute-angled sprockets is 8 mm, the chips from

their impacts converge, which reduces the surface roughness.

The mills were tested on concrete tiles-samples of size 25 × 440 cm with compressive strength of 10, 15, 20 and

30 MPa. The best effect was manifested for medium and high strength concrete: 15–20 MPa (Aggregate - Marble and

Limestone) and 30 - 35 MPa (Aggregate - Granite). Concretes of low strength (10 - 15 MPa) are processed less

efficiently due to punctures of coarse aggregate, the formation of shells and the collapse of the edges of the grooves.

At the same time, sprocket wear was assessed as a percentage by weight. The wear of the sprockets from various types

of steel was: R18 steel - 0.16%, steel 40Kh - 0.16%, steel ShKh-15- 0.17%, VK8 plates - 0.01%.

Treatability of concrete and reinforced concrete depends mostly on abrasive and strength characteristics of its

components – fine and coarse filler that might make up to 80% of the total volume of concrete. Therefore, the choice

of mechanical treatment parameters depends on treatability of rock used for production of concrete fillers. Based on

the test results, preliminary conclusions were made concerning applicability of inertial tool stars for treatment of

concrete featuring different physical and mechanical properties of fillers (Table 1).

Table 1. Characterization of inertia tool materials for matured concrete treatment

Type of Concrete and Reinforced Concrete Recommended Steel Grade

Heavy concrete based on silicate rock fillers with compression strength of the input

rock of 250 to 450 MPa (Granites, Basalts, Sandstones) VK8, 40Kh

Heavy concrete based on carbonate rock fillers with compression strength of input

rock of 150 to 250 MPa (Tight Limestones) R18, 40Kh

Heavy concrete based on carbonate rock fillers with compression strength of input

rock up to 150 MPa (Marbles) 40Kh, ShKh-15

4. Conclusion

Taking into account the importance and complexity of façade finishing jobs in the total scope of construction

operation, selection of the best labor practices for their performance is a highly relevant matter. The straightforward

segmented work flow is widely used in construction. The practical benefit of splitting the complex construction

process into simple processes or particular operations is that it allows to speed up jobs and improve labor productivity

thanks to workers’ specialization focused on the performance of particular repetitive operations. Compiling teams and

crews of workers having appropriate qualifications, utilization of high-performance tools and appliances, and, first of

all, rational workplace management involving splitting of the spread of works into divisions and plots allow reducing

labor costs and achieving a better quality of works carried out.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1005

The types and location of surfaces subject to stucco jobs render considerable influence on operational efficiency of

teams. One of the opportunities for increasing labor productivity and efficiency during outdoor stucco operations is

improving powered tools utilized to treat outer wall surfaces of buildings. The technical and economic characteristics

of the process of mechanical treatment of building wall materials (stucco, brick, concrete, reinforced concrete) depend

on the correct choice of tool modes and parameters. Parameters should be selected with regard to specific conditions

based on physical and mechanical characteristics of wall material and tool. In the construction site setting, it is useful

to get to know the following necessary physical and mechanical properties of materials to be treated: material’s

compression strength; composition of coarse and fine filler for concrete and reinforced concrete; presence, diameter

and layout of bars when cast-in-situ structures are treated to a depth of 5 - 8 mm. The laboratory tests performed have

allowed selecting optimal modes of tool operation for heavy concretes and compare efficiency of different steels for

the fabrication of mill stars. It is necessary to continue investigating into selection of the mode and parameters of tools

for mechanical treatment of surfaces as applied to special types of concrete: polymer concretes based on silicate and

carbonate fillers, concretes based on porous fillers (Tuff, Pumice, Slag, Keramzit). The Cost-efficiency of steels and

alloys applied should be studied, too.

5. Conflicts of Interest

The authors declare no conflict of interest.

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monolithic structures in winter” BST Bulletin of Construction Equipment 4. (2018).

[2] Ershov M. N., Lapidus A. A., Telichenko V. I. Tekhnologicheskie protsessy v stroitel'stve (The technological processes in

construction), Moscow, ASV Publ., 2016. (In Russian).

[3] Gumerova, Eliza, Olga Gamayunova, and Roman Gorshkov. “Choosing the Appropriate Way of Plastering Works for

Transportation and Construction Facilities.” IOP Conference Series: Earth and Environmental Science 90 (October 2017):

012185. doi:10.1088/1755-1315/90/1/012185.

[4] Vatin, Nikolay Ivanovich, and Olga Sergeevna Gamayunova. “Using Plastering Machines to Improve the Efficiency of

Finishing Works.” Applied Mechanics and Materials 635–637 (September 2014): 2049–2053.

doi:10.4028/www.scientific.net/amm.635-637.2049.

[5] Makovetskaya, Elena, Antonina Deniskina, Egor Krylov, and Fatima Urumova. “Organizational Optimization of Construction

Processes by Virtue of Robotization.” Edited by A. Zheltenkov. E3S Web of Conferences 91 (2019): 02036.

doi:10.1051/e3sconf/20199102036.

[6] Bock, Thomas, Natalia Buzalo, and Alexey Bulgakov. “Mathematical Description and Optimization of Robot Control for

Plastering Works.” 2018 International Multi-Conference on Industrial Engineering and Modern Technologies (FarEastCon)

(October 2018). doi:10.1109/fareastcon.2018.8602717.

[7] Shreeranga B., Nishchith H., A. Kumar U.R., Ramith R., Naveen M.V., Jishnumohan D. Nair. “Design and Fabrication of Wall

Plastering Machine”, Journal of Mechanical Engineering and Automation, 7(5): (2017); 159-163.

doi:10.5923/j.jmea.20170705.07.

[8] Ankush, N.A., Laukik, P.R. “Design of automatic wall plastering machine” International journal of engineering sciences &

research technology, (2017): 6 (3). 543-555. DOI: 10.5281/zenodo.439237.

[9] Tishkin, D. D., and K. I. Barbolin. “To the Issue of Improving the Durability of the Plaster Facades of Buildings.” Bulletin of

Civil Engineers 14, no. 6 (2017): 135–139. doi:10.23968/1999-5571-2017-14-6-135-139.

[10] Pakhomova, L.A., Chernyshova, A.M. “Organizational and Technological Solutions for the Application of Leveling Finishing

Coatings”, Science Prospects, 2:101, (2018): 62-71.

[11] Adams, Thomas, Anya Vollpracht, Johannes Haufe, Linda Hildebrand, and Sigrid Brell-Cokcan. “Ultra-Lightweight Foamed

Concrete for an Automated Facade Application.” Magazine of Concrete Research 71, no. 8 (April 2019): 424–436.

doi:10.1680/jmacr.18.00272.

[12] Gerek, Ibrahim Halil, Ercan Erdis, Gulgun Mistikoglu, and Mumtaz A. Usmen. “Evaluation of Plastering Crew Performance

in Building Projects Using Data Envelopment Analysis.” Technological and Economic Development of Economy 22, no. 6

(September 25, 2014): 926–940. doi:10.3846/20294913.2014.909903.

[13] Idiake, John Ebhohimen, and Mbamali Ikemefuna. "Improving Labour Performance in the Management of Wall Plastering

Activity for One Storey Buildings in Abuja, Nigeria." Journal of Economics and Sustainable Development (2004).

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[14] Bokor, Orsolya, Laura Florez, Allan Osborne, and Barry J. Gledson. “Overview of Construction Simulation Approaches to

Model Construction Processes.” Organization, Technology and Management in Construction: An International Journal 11, no.

1 (March 1, 2019): 1853–1861. doi:10.2478/otmcj-2018-0018.

[15] Lapidus, A.A., Tolstova, K.S., Topchy, D.V. “The Formation of Groups of Parameters Affecting the Evaluation Criteria of

Permissibility of Combining Streams in the Manufacture of Building Finishing Work in Residential Buildings”, Science and

business: development ways, 2018, 6 (84): 18-22

[16] Zhadanovsky, Boris. “Mechanical Processing of Concrete and Reinforced Concrete with Diamond Tool.” Edited by A.

Mottaeva and B. Melović. MATEC Web of Conferences 193 (2018): 03013. doi:10.1051/matecconf/201819303013.

[17] Grankina, D.V., Troitskiy, V.M., Vasilieva, D.K., Matycin A.V. “Theoretical bases of processing of building structures with

free abrasives”, Engineering journal of Don, 1 No. 52, (2019).

[18] Oleynik, P.P., Brodsky, V.I., and Cherednichenko, N.D., “Theory, Methods, and Forms of Organization of Building

Production”, part 1, publishing house Misi-Mgsu, 2019: 335.

[19] Vainshtein, M.S., Zhadanovsky, B.V., Sinenko, S.A., et al. “Evaluation of the effectiveness of organizational and

technological solutions when choosing means of mechanization of construction and installation works”, Scientific Review, vol

13, (2015):156-159.

[20] Opanasyuk, I.L., Opanasyuk, L.G., Reutsky, I.A., and Paytra, A.P. “Reserves for increasing the efficiency of finishing work in

the construction of residential and public buildings”, Bulletin of the Belarusian-Russian University, 3 No. 40, (2013): 82-91.

Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

1007

Microstructural and Compressive Strength Analysis for Cement

Mortar with Industrial Waste Materials

Zahraa Fakhri Jawad a, Rusul Jaber Ghayyib

a, Awham Jumah Salman

a*

a Al-Furat Al-Awsat Technical University, Najaf, Kufa, Iraq.

Received 06 December 2019; Accepted 02 March 2020

Abstract

Cement production uses large quantities of natural resources and contributes to the release of CO2. In order to treat the

environmental effects related to cement manufacturing, there is a need to improve alternative binders to make concrete.

Accordingly, extensive study is ongoing into the utilization of cement replacements, using many waste materials and

industrial. This paper introduces the results of experimental investigations upon the mortar study with the partial cement

replacement. Fly ash, silica fume and glass powder were used as a partial replacement, and cement was replaced by 0%,

1%, 1.5%, 3% and 5% of each replacement by the weight. Compressive strength test was conducted upon specimens at

the age of 7 and 28 days. Microstructural characteristic of the modified mortar was done through the scanning electron

microscope (SEM) vision, and X-ray diffraction (XRD) analysis was carried out for mixes with different replacements.

The tests results were compared with the control mix. The results manifested that all replacements present the

development of strength; this improvement was less in the early ages and raised at the higher ages in comparison with the

control specimens. Microstructural analysis showed the formation of hydration compounds in mortar paste for each

replacement. This study concluded that the strength significantly improved by adding 5% of silica fume compared with

fly ash and glass powder.

Keywords: Cement Replacements; Fly Ash; Silica Fume; Glass Waste; Recycling Materials; Compressive Strength.

1. Introduction

Production of cement implicates a high consumption of energy and therefore is responsible for almost (7%) of the

world’s (CO2) emission. It’s well known that (CO2) is the main contributor to the greenhouse influence and

subsequently being responsible for the global warming of the earth. Thus, research upon the usage of by-product

cementing materials, like silica fume, metakaolin, fly ash, waste glass and rice husk ash in place of cement has been

increased in concrete technology [1].

Currently, researches on sustainable development on concrete have been carried out on the following aspects:

extension of “concrete structure and development of low-carbon concrete material and structure [2]. Contemporary

mixed cement types also use pozzolan as a cement replacement material or a mineral additive that is inter-ground or

mixed with Portland cement. Pozzolans are defined as “Siliceous or Siliceous-Aluminous materials that have a slight

or no cementitious effect, but due to their fine separated shapes and with the existence of moisture. They are

chemically reacting with the calcium hydroxide under normal temperature to produce compounds having cementitious

characteristics [3]. Krishnaraj et al. (2017) explored the influence of fly ash on the durability characteristics and

compressive strength of the mixed cement mortar. The compressive strength test results and the (SEM) analysis

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091524

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1008

depicted that incorporating fly ash in mortar accelerates the (C-S-H) gel development via using the content of (C-H),

causing better durable and mechanical properties of mortar [4]. Khalaf et al. (2018) investigated the influence of the

local fly ash on the cement mortar characteristics. This study found that the compressive strength values were declined

for all the fly ash mixes at the early ages, 3 days. However, the compressive strength was improved remarkably at the

later ages 28 days, [5]. Yerramala et al. (2012) evaluated the strength characteristics of the fly ash mortars; results

revealed that at the early age of all replacements of fly ash, the strength reduced to the ordinary mortar .Nevertheless,

after 28 days and beyond, mortars prepared with the replacement of fly ash up to 15% caused a higher strength than

the ordinary mortar [1].

Bhandari and Tajne (2013) investigated the mechanical characteristics of mortar blocks possessing different levels

of replacements of fine and coarse waste glass with fine aggregate, the results of the test showed that the fine

aggregate replacement via fine glass has an essential influence upon the mortar blocks compressive strength in

comparison with the control sample due to the pozzolanic nature of (FG). The results elucidated the pozzolanic

reactivity of such waste and the open potentials for the use of this material in mortars [6]. Aseel et al. (2016) studied

the mechanical characteristics and thermal conductivity of mortar cement produced from glass waste at different

values of glass to cement. Ultimately glass waste can be used as an environmentally and cost-effective cement

substitute in mortar cement production [7].

Amudhavalli et al. (2012) considered the influence of the partial cement replacement by (0, 5, 10, 15 and 20%)

silica fume. The results of the test depicted that the silica fume usage in concrete enhanced the concrete performance

in strength and durability features [8]. Sasiekalaa et al. (2012) studied the mortar compressive strength for Ferro

cement containing ternary mixes of silica fume, Portland cement, and superplasticizer as a water reducing agent. The

results revealed that the compressive strength cement mortar enhanced with silica fume content [9]. Tayeh et al.

(2019) confirmed that the use of glass powder in mortar explored the strength properties of mortar and its sulphate

resistance, moreover, the results indicated that replacing glass powder in the control mixture increased the unit weight

by 29% at 60 days [10]. Czapik and Mateusz et al. (2018) evaluated the effect of silica fume on the properties of

cement mortar. This study confirmed the improvement in compressive strength achieved when added silica fume to

cement paste [11]. Ortega et al. (2018) studied the long-term effects 400 days of the addition of waste glass powder on

the microstructure and service properties of mortars, which incorporate up to 20% of this addition as clinker

replacement. And that showed good service properties until 400 days, similar to or even better than those made with

ordinary Portland cement without additions, with the added value to contributing to sustainability [12].

This research was conducted to study the effect of using some sustainable materials which have pozzolanic

properties such as silica fume, fly ash and fine glass powder which would be introduced to cement mortar technologies

with different content. Therefore, research themes will focus on befits and determinants of the implementation of these

replacements in cement and concrete technology. Also, the optimum content for each type would be recognized.

2. Experimental Method

2.1. Materials

In the current research, the ordinary Portland cement (Type I) confirmed to the Iraqi Standard Specifications No.

5 [13] was used and the chemical compositions of cement shown in Table 1. Silica fume, fly ash and glass powder

were used as a partial cement replacement, Figure 1 indicates the particle size analysis for each of them. The silica

fume used in this experiment contained (94.6%) SiO2, the chemical compositions of fly ash and glass powder were

analyzed by an (XRF) microprobe analyzer and depicted in Table 2. Glenium 54 (G54) high range water reducing

admixture type F was introduced into the whole blends. Natural sand of river was utilized with the portion of sand that

passes through 850 µm sieve according to Iraqi Standard specification no 2080 [14].

Table 1. Chemical composition of cement

Oxide Content, %

CaO

SiO2

Al2O3

Fe2O3

MgO

K2O

Na2O

SO3

66.11

21.93

4.98

3.10

2.0

0.75

0.35

2.25

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1009

(a)

(b)

(c)

Figure 1. Particle Size Analysis for: (a) Silica Fume (b) Fly Ash and (c) Glass Powder

Table 2. Chemical compositions of fly ash and glass powder

SiO2 MgO K2O CaO Al2O3 Cl Fe2O3 Na2O SO3 CuO ZnO

Fly ash 88.6 3.21 3.26 2 0.75 3.26 1.72 - - - -

Glass 72.5 3.72 - 10.2 0.7 - 0.31 9.9 0.244 0.227 0.133

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2.2. Mixture Details

In this research, thirteen mixtures of mortar consist of (cement, sand, water, superplasticizer and replacements) to

prepare the samples. Constant water/cement ratio (𝑤/𝑐) = 0.45 and superplasticizer Percentage=1.25% by the

cement weight were used for all mixture. The mortar mixtures details illustrated in Table 3.

Table 3. Mortar mixes

Mix Symbol Cement

(g)

Sand

(g)

Fly ash

(g)

Silica fume

(g)

Glass powder

(g)

Control 500 1375 - - -

1F 495 1375 5 - -

1.5F 492.5 1375 7.5 - -

3F 485 1375 15 - -

5F 475 1375 25 - -

1SF 495 1375 - 5 -

1.5 SF 492.5 1375 - 7.5 -

3SF 485 1375 - 15 -

5SF 475 1375 - 25 -

1GP 495 1375 - - 5

1.5GP 492.5 1375 - - 7.5

3GP 485 1375 - - 15

5GP 475 1375 - - 25

F: is the mixes with Fly Ash. SF: is the mixes with Silica Fume. GP is the mixes with Glass Powder

2.3. Compressive Strength Measurements

According to ASTM C109, (50 × 50 × 50) mm mortar cubes were cast using the mix proportion shown in Table 3,

w/c ratio of 0.45. Specimens were cast, and throughout the moulding, and the cubes were mechanically shaken.

Beyond 24 hours, these specimens were de-moulded and subjected to curing in distilled water with two groups the

first group cured for 7 and the second group were cured for 28 days. And, beyond curing, such specimens were

examined for the compressive strength utilizing a compression testing machine. The tests carried out according to BS

1881-Part 101 [15] on three specimens, and the average values of compressive strength were determined. Compression

test was done using BESMAK testing equipment.

3. Results and Discussion

3.1. Results of Compressive Strength

Figures 2 to 4 demonstrate the effect of replacement type on the compressive strength of mortar's specimens after 7

and 28 days of curing. In Figure 2, the results show clearly the mortar mixes, where the silica fume was used as partial

replacement, in this figure the compressive strength increased gradually with the increase of silica fume contents, it

raised from 27 MPa for control mortar to 41 MPa with the increasing percentage of (5%). The behaviour of improving

the compressive strength with the increase of silica fume in mortar mixtures is due to that if the silica fume introduced,

the cement hydration rate raises at the initial hours owing to release of (OH) ions and alkalis into the pore fluid. Also,

additional SiO2 will react with CH to produce more CSH gel [16]. This further formation of CSH gel in the cement

paste enhanced the mortar strength properties and gave a more dense structure due to the filler effect of silica fume.

These results agreed with Al Ghaban et al. (2018) [17] and Rostami and Behfarnia (2017) [18].

In Figure 3, the results indicate that the mixes with fly ash have no appreciable early strength. However, those

mixes have gained the strength more than the target strength in later days this phenomenon of gaining strength with

late age when using fly ash as partial replacement of cement was demonstrated by Yerramala et al. (2012) [1] and

İlhami et al. (2017) [19] and other researchers. The compressive strength was increased from (27) MPa for control

mortar to (39.5) MPa for mixes with (5%) fly ash at 28 days. This development in compressive strength might be

attributed to the chemical reactivity of fly ash with CH to form CSH gel and thus gain binding property in cement

paste according to Krishnaraj et al. (2017) study [4]. In Figure 4, the results reveal, similarly to those of mixes with

silica fume and fly ash, that the powder of waste glass incorporation enhanced the cement mortars mechanical

strength. Due to the pozzolans in the glass powder composition which would react with the available (CH), chiefly

making (CSH) similar to that formed in the hydration reactions of cement, the results of an approach obtained by

Sofiane et al. (2019) [20] and Tamanna et al. (2016) [21]. Thus, confirming that waste glass powder can further

contribute to sustainability in construction. Comparing the results of compressive strength yielded that silica fume

showed more significant improvement in strength by analogy with fly ash and glass powder respectively, this

distinction for silica fume can be attributed to its high silica content with the high pozzolanic effect. However, it can

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1011

be concluded that all the studied replacements can be considered as a viable replacement for cement and is thus an

economical construction material.

Figure 2. Results of the compressive strength test for the concrete mixes with silica fume as cement partial replacement

Figure 3. Results of the compressive strength test for the concrete mixes with fly ash as cement partial replacement

Figure 4. Results of the compressive strength test for the concrete mixes with glass powder as partial cement replacement

22

32

33

.2

35

36

27

38

40

41

.5

41

0 % 1 % 1 . 5 0 % 3 % 5 %

Co

mp

ress

ive

Str

ength

MP

a

Replacement Content%

7 day 28 day

22

24

25

27

29

.2

27

30

33

.7

35

39

.5

0 % 1 % 1 . 5 0 % 3 % 5 %

Co

mp

ress

ive

Str

ength

MP

a

Replacement Content%

7 day 28 day

22

25

26

31

37

27

31

34

.3

37

40

0 % 1 % 1 . 5 0 % 3 % 5 %

Co

mp

ress

ive

Str

ength

MP

a

Replacement Content %

7 day 28 day

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1012

3.2. Results of SEM Analysis

From the SEM analysis of blended cement mortar samples, it is evident that no clear particles of silica fume, fly

ash or fine glass powder are seen; hence, it proved that all the pozzolanic materials contribute to the hydration process.

Reference sample microstructure is shown in Figure 5(a). When these displays were examined, it was observed that

the (C-S-H) gels started to form in the cement pastes, Figure 5(b) views that the formation of (C-S-H) gels increased

in samples with the silica fume. In Figure 5(c, d) for the (SEM) images of mixes with fly ash or fine glass powder, the

(C-S-H) gels observed to be more significant and that they covered most of the internal structure. These findings

confirm by Ana Mafalda Matos et al. carried out a study concerning the microstructure of mortar containing waste

glass powder [22].

The SEM images of the hardened cement paste show that the distribution of (CSH) nearly increased compared to

the reference mix due to the replacement of cement by pozzolanic replacements (silica fume, fly ash and glass

powder). In the mixtures, the development in the (CSH) occurred as a result of the reaction between (CH) and

pozzolanic replacing materials to produce extra (CSH) and ettringite. This mechanism was accordance with the

mechanism proposed by Henry Limantono et al. (2016), explaining excellent hydration between cement, glass powder,

and silica fume resulted in a massive layer of paste due to the pozzolanic effect [23]. Also, the filler action of presence

the replacement particles which play the primary role in increasing the density of modified mortar compared with

reference mortar. This increasing explains the reason behind the strength development when using these materials as

partial cement replacement which can be analyzed and explained based on the growth of hydration products in the

microstructure of mortar mixes [23, 24].

(a) (b)

(c) (d)

Figure 5. SEM photomicrograph for mortar mix with: (a) Reference Mix, (b) Silica Fume Mix, (c) Fly Ash

Mix, (d) Glass Powder Mix

CS

H

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1013

3.3. Results of XRD Analysis

(XRD) analysis performed to investigate the replacements of cement reactivity in the mortar after a curing time of

28 days. Figures 6 to 9 depict the XRD with angular range (2θ = 0 - 60) degree for reference mix and mixes with silica

fume, fly ash and glass powder, correspondingly. The primary phases present in the Portland cement after the

hydration process were calcium silicate hydrate gel (C-S-H), Calcium Hydroxide (CaOH2), calcium aluminate

(CaAl2O4), calcium silicate (CaSiO3) and ettringite phases. In all XRD spectra, the same three main types of the

observed peaks were Calcite (CaCO3) with a hexagonal crystal system [26], Quartz (SiO2) with a hexagonal crystal

system [27], Portlandite (CaOH2). Among them, SiO2 mainly came from the siliceous sand, and CaCO2 produced from

the addition mixed with the cement during its production process and carbonation of cement hydration products, and

Ca(OH)2 was formed from the hydration of cement clinker. This result confirms the finding by Huang and Zhao

(2019), who studied the correlation between strength and durability of mortar with fly ash [28]. The hydration process

in a mortar with mineral admixture is very complicated as compared with the hydration of Portland cement. Therefore,

all (XRD) charts in this research present the calcium silicate hydration (C-S-H) unmanageable to distinguish the glassy

phase (lack in crystallinity). Calcium hydroxide (CH) is also difficult to differentiate due to the pozzolanic reaction,

which forms calcium silicate hydration (C-S-H) that leads to depleting unknown amount of calcium hydroxide (CH)

this was accordance with the results proposed by Sang-Hwa Junget.al, explaining the phases corresponding to

hydration of mortar with fly ash and RHA [29]. The results indicate that calcium hydroxide appears at very low peak

intensity, as shown in the previous figures because of the consuming of calcium hydroxide (CH) by the pozzolanic

reaction that indicates the pozzolanic reactivity of the partial replacement materials, and that is supported by the

compressive strength and scanning electron microscope results [30, 31].

Figure 6. The characterization measurement (XRD) pattern of reference mortar mix

Figure 7. The characterization measurement (XRD) pattern of mortar mix with silica fume

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1014

Figure 8. The characterization measurement (XRD) pattern of mortar mix with fly ash

Figure 9. The characterization measurement (XRD) pattern of mortar mix with glass powder

4. Conclusions

The following conclusions are drawn from the obtained experimental study:

Silica fume, fly ash and glass powder may be considered as a critical cement replacement admixtures in

building up the physical strength of hardened cement paste.

Silica fumes considered to be a proper admixture added to improve early and later strengths of mortar,

furthermore present significant improvements to the compressive strength compared with fly ash and glass

powder.

The presented results obtained using (SEM) and (XRD) manifested the reasonable chance for observing the

variations in the maturing cement pastes microstructure depending upon the differences in the morphology of

the main phases arising throughout the cement paste hydration which help in predicting the properties of

cement mortars.

In the (SEM) micrographs, the dominant phases present in microstructure were Portlandite (CH), (CSH), and

ettringite, which considered the evidence of hydration process, it was clear that the microstructural behaviour of

mortar influences the strength characteristics of the mix.

The XRD results elucidated that the addition of silica fume, fly ash and glass powder effects on the hydration

product quantity and distribution.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

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5. Conflicts of Interest

The authors declare no conflict of interest.

6. References

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[3] Sobolev, Konstantin. “Sustainable Development of the Cement Industry and Blended Cements to Meet Ecological Challenges.”

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[4] L. Krishnaraj, Yeddula Bharath Simha Reddy, N. Madhusudhan and P.T. Ravichandran, “Effect Of Energetically Modified Fly

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[5] Khalaf, Ali Abdulhasan, Fadhil Kamil Idan, and Kadhim Zuboon Nasser. "Effect the Local Fly Ash on Cement Mortar

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[7] Aseel Basim Al-Zubaidi, Ahmed A. Al-Tabbakh, “Recycling Glass Powder and its use as Cement Mortar applications”,

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[12] Ortega, José, Viviana Letelier, Carlos Solas, Marina Miró, Giacomo Moriconi, Miguel Climent, and Isidro Sánchez. “Influence

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(March 16, 2018): 842. doi:10.3390/su10030842.

[13] Iraqi Standard Specifications No. 5, “Portland Cement”, Central Organization for Standardization and Quality Control, Iraq,

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[15] BS 1881-116, “Testing concrete — Part 116: Method for Determination of Compressive Strength of Concrete Cubes”, British

Standard, December 2003.

[16] Pranab Chakraborty, “Investigation on Flexural Strength of High Strength Silica Fume Concrete”, International Research

Journal of Engineering and Technology, Vol. 4, (2016):1722-1726.

[17] Al Ghaban, Ahmed, Aseel Al Zubaidi, and Zahraa Jawad. “Study The Effect of Micro CaCO3 and SiO2 and Their Mixture on

Properties of High Strength Concrete.” Engineering and Technology Journal 36, no. 10A (October 25, 2018).

doi:10.30684/etj.36.10a.2.

[18] Rostami, M., and K. Behfarnia. “The Effect of Silica Fume on Durability of Alkali Activated Slag Concrete.” Construction and

Building Materials 134 (March 2017): 262–268. doi:10.1016/j.conbuildmat.2016.12.072.

[19] İlhami Demir, Selahattin Güzelkücük, Özer Sevim, Ahmet Filazi, and Çağri Göktuğ Şengül, “Examination of Microstructure

of Fly Ash in Cement Mortar”, Conference: International Conference on Engineering and Natural Science (ICENS), At

Lisbon, Portugal 2017.

[20] Sofiane Saggai, Saci Dahmani, Marwa Boulifa, Adel Debbabi, “Waste Glass Powder in mortar: technical and environmental

effects”, Conférence Internationale surles Matériaux, 2019.

[21] Tamanna, Nafisa, Norsuzailina Mohamed Sutan, Rabin Tuladhar, Delsye Teo Ching Lee, and Ibrahim Yakub. “Pozzolanic

Properties Of Glass Powder In Cement Paste.” Journal of Civil Engineering, Science and Technology 7, no. 2 (September 30,

2016): 75–81. doi:10.33736/jcest.307.2016.

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[22] Matos, Ana Mafalda, and Joana Sousa-Coutinho. “Waste Glass Powder in Cement: Macro and Micro Scale Study.” Advances

in Cement Research 28, no. 7 (July 2016): 423–432. doi:10.1680/jadcr.14.00025.

[23] Limantono, Henry, Januarti Jaya Ekaputri, and Tri Eddy Susanto. “Effect of Silica Fume and Glass Powder on High-Strength

Paste.” Key Engineering Materials 673 (January 2016): 37–46. doi:10.4028/www.scientific.net/kem.673.37.

[24] Vikas Srivastava, Shivacharan Singh. “Glass Waste in Concrete: Effect on Workability and Compressive Strength.”

International Journal of Innovative Research in Science, Engineering and Technology 4, no. 9 (September 15, 2015): 8142–

8150. doi:10.15680/ijirset.2015.0409018.

[25] Franus, Wojciech, Rafal Panek, and Magdalena Wdowin. “SEM Investigation of Microstructures in Hydration Products of

Portland Cement.” 2nd International Multidisciplinary Microscopy and Microanalysis Congress (2015): 105–112.

doi:10.1007/978-3-319-16919-4_14.

[26] Tomaszewski, P. E. "Golden book of phase transitions." Wroclaw 1 (2002): 1-123.

[27] Kern, A., and W. Eysel. "Mineralogisch-Petrograph." Inst., Univ. Heidelberg, Germanny, ICDD Grant-in-Aid (1993).

[28] Huang, Qian, and Liang Zhao. “Correlation between Compressive Strengths and Water Absorption of Fly Ash Cement Mortar

Immersed in Water.” Archives of Civil Engineering 65, no. 3 (September 1, 2019): 141–152. doi:10.2478/ace-2019-0040.

[29] Jung, Sang-Hwa, Velu Saraswathy, Subbiah Karthick, Palanivel Kathirvel, and Seung-Jun Kwon. “Microstructure

Characteristics of Fly Ash Concrete with Rice Husk Ash and Lime Stone Powder.” International Journal of Concrete

Structures and Materials 12, no. 1 (February 21, 2018). doi:10.1186/s40069-018-0257-4.

[30] Dhapekar, N. K., A. S. Majumdar, and P. K. Gupta. "Study of phase composition of Ordinary Portland Cement concrete using

X-Ray diffraction." International Journal of Scientific and Engineering Research 6, no. 11 (2015).

[31] Matsushita, Tetsuro, Hiroshi Hirao, Ippei Maruyama, and Takafumi Noguchi. “Quantitative Analysis of Cement Hydration by

XRD/Rietveld Analysis.” Journal of Structural and Construction Engineering (Transactions of AIJ) 73, no. 623 (2008): 1–8.

doi:10.3130/aijs.73.1.

Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

1017

Improving the Aging Resistance of Asphalt by Addition of

Polyethylene and Sulphur

Maria Iqbal a*

, Arshad Hussain a, Afaq Khattak

b, Kamran Ahmad

a

a Department of Transportation, School of Civil and Environmental Engineering, National University of Science and Technology,

Islamabad, Pakistan.

b International Islamic University, Islamabad, Pakistan.

Received 18 November 2019; Accepted 09 March 2020

Abstract

With the increase in demand of flexible pavements, due to their various advantages over rigid pavements, there is a need

to improve the aging properties of the bitumen in order to enhance its resistance against different types of distresses such

as rutting, fatigue cracking. This research focus on the use of one polymeric additive Polyethylene (PE) and one non

polymeric additive Sulphur (S) to enhance the aging resistance of asphalt. These modifiers are evaluated for their effect

on the aging mechanism in comparison with the unmodified bitumen. Aging of the original and modified bitumen is

realized by the Rolling Thin Film Oven (RTFO) and Pressure Aging Vessel (PAV). Physical properties of the aged and

unaged asphalt binders are evaluated through empirical testing like penetration, ductility and softening point test.

Optimum content of the modifiers is obtained by comparing the results of conventional properties before and after aging.

Fourier Transformed Infrared Spectroscopy (FTIR) and Scanning Electron Microscope (SEM) are performed to bring out

the chemical and morphological changes in the modified binder. Rheological properties of modified asphalt are evaluated

with the help of a Dynamic Shear Rheometer (DSR). Results indicate improvement in physical properties of the modified

asphalt even after the aging. Penetration index increased which shows less temperature susceptibility of the modified

binders. Carbonyl and sulfoxide index are used as aging indicators which shows reduction in case of modified samples.

Decrease in the sulfoxide and carbonyl index indicates better oxidation resistance of the modified samples.

Morphological analysis proves good compatibility of the modifiers with asphalt binders. DSR results indicate improved

viscoelastic properties of the modified binders. Hence it can be concluded that Polyethylene and Sulphur are good

options to improve the aging resistance of asphalt in terms of their cost effectiveness and environment friendly nature.

Keywords: Bitumen Modification; Aging; Binder Aging; Polyethylene; Sulphur.

1. Introduction

Pakistan is a developing country where roads and highways are major source of transportation. Almost 95% of the

population and freight movement is served by its major highway [1]. Due to rapid increase in transportation through

roads, demand for bituminous pavement has also been increased as flexible pavements are more economical and

provide smooth riding quality [2, 3]. Despite the lower construction cost, the life cycle cost of the flexible pavement is

higher than the rigid pavement due to its higher maintenance and rehabilitation cost [3].

In flexible pavements, mostly, failure of the highway occurs due to the non-structural rutting or different types of

cracking. Non-structural rutting occurs due to the poor asphalt mixture properties, heavy traffic loads, or due to high

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091525

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1018

temperatures. Aging is one of the prime causes of early failure of asphalt pavement [4-6]. Aging is defined as the

oxidation of light components which causes the stiffening of bitumen during construction and service phase [5, 7]. It

affects the chemical and rheological properties of bitumen, causing it to fail before the estimated service life. Asphalt

aging affects the pavement flexibility negatively after many years of service life in field [8]. The process of aging is

greatly dependent on the chemical composition of bituminous mixture [4]

A lot of research has been carried out to enhance the aging resistance of bitumen. Various types of polymers like

Styrene Butadiene Styrene (SBS) [9], polyethylene [9], polypropylene [9] etc., have been used in asphalt pavements

previously. Studies show that these polymer modifiers bring significant improvements in the mechanical, rheological

and physical properties of the bitumen. They also found that the type of chemical additive and the aging temperature

considerably affect the recovery behavior of asphalt binder [10].

Modification of asphalt binder can be done by adding different percentages of elastomers up to 7%. Soft

modifications contain a polymer content of up to 3% while medium modifications have polymer content of about

4.2%. Hard modifications have polymer content higher than 5% [11]. 3% LLDPE improved the mixture stiffness at

40°C as well as the rut depth was also improved measured by wheel tracking test [12]. PE modified bitumen is also

tested to tackle the problem of phase separation and storage stability [13]. 3% polyethylene content was suggested to

be the optimum content. Polyethylene content of 5% or more was regarded as not applicable because the polyethylene

modified bitumen became unworkable at these percentages due to the high values of rotational viscosity [14].

Hamburg wheel track test, resilient modulus, dynamic creep and indirect tensile test were conducted on the PE

modified concrete mixture. The analysis showed that the PE modified blends gave better performance results than the

conventional asphalt mixtures. Temperature susceptibility and rutting resistance was improved. The authors suggested

a PE content of 5% to be used for better performance of asphalt mixtures [15].

SBS and sulfur were tested to enhance the storage stability of TR modified bitumen. Results suggested that the

Penetration Temperature Susceptibility (PTS) of bitumen was largely decreased. Storage stability improved to higher

extent and the Marshall stability and plastic deformation resistance was also enhanced with the addition of sulphur

[16]. Addition of Sulphur decreases the viscosity of bitumen. Lower viscosity will lead to decreased mixing and

compaction temperature making it more economical and energy efficient [17]. 2% sulfur by weight of asphalt binder

at 140°C mixed for 30 minutes results in the best homogenous modification of VG30 binder [18].

Calculation of aging index is one of the most realistic approaches to observe the impact of aging on various

characteristics of bitumen. Aging index of asphalt can be defined as the ratio of some property of asphalt after aging to

the same property of asphalt before aging as given in Equation 1 [19].

Aging index =binder property after aging

same property before aging (1)

The aim of this study is to improve the rheological properties of asphalt in terms of aging by using cheap and

environment friendly modifiers. Polyethylene which is the main source of plastic is left undisposed, causing the huge

waste pollution. Similarly, now days, almost all elemental Sulphur obtained is the by-product of gas and petroleum.

Therefore the incorporation of abundantly existing materials into pavement industry is a technique which is both cost

effective and environment friendly .Moreover, the addition of the modifiers reduces the cost of the pavement as well

as increases the service life of flexible pavements by decreasing the amount of distresses whose major cause is the

phenomena of aging in asphalt.

2. Research Methodology

Methodology adopted in conducting this research is summarized in the chart below;

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1019

Figure 1. Flow chart of experimental work

2.1. Material Selection

Penetration grade of 60/70 bitumen supplied by PARCO sales office Rawalpindi was used as base bitumen. Table

1 shows the basic properties of base binder. The purpose of selecting grade 60/70 is that it is typically used in Pakistan

and is appropriate for colder to Intermediate temperature regions. Polyethylene (PE) and Sulphur (S) were used as

modifiers. Polyethylene used in this research is in liquid form while Sulphur is used in powder form. Material are

selected as per the availability and cost efficiency of the material. Physical properties of polyethylene and Sulphur are

listed in Tables 2 and 3 respectively.

Table 1. Physical properties of 60/70 asphalt

Property AASHTO Designation Result AASHTO Specification

Penetration T49 66 60-70

Softening Point T53 49 49/56

Ductility T51 117 100 cm

Specific Gravity

1.02 1.04 Max

Viscosity at 135.5°C T316 450cP ≤3 Pa.s

Table 2. Physical properties of polyethylene

Properties Result

Chemical Formula (C2H4)n

Melting Point 115-135 °C

Density 0.88–0.96 g/cm3

Table 3. Physical properties of sulphur

Properties Result

Appearance Yellow Crystalline Solid

Specific Gravity 1.92

Melting Point 120°C

Experimental Program

Testing

Conventional Testing

Penetration

Softening Point

Ductility

Advance Testing

FTIR

SEM

DSR

Material

60/70 PARCO Bituemn

Polyethylene

Sulphur

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1020

2.2. Preparation of Modified Binder

Samples were prepared by adding three percentages 2, 3 and 5% by weight of polyethylene (PE) and Sulphur (S)

into 60/70 base binder. 500g of bitumen for each percentage of each modifier was heated until it turned into liquid.

High rate shear mixer was used for the mixing of modifiers into base bitumen at 140°C and 1200rpm. The mixing was

continued for about 30 minutes so that P.E and S can be completely dissolved in asphalt binder.

2.3. Aging of the Bitumen

Rolling thin film oven (RTFO) was used to feign the effect of short term aging on neat and modified binders.

Asphalt in RTFO is aged by heating and blowing of hot air at 163ºC for 80 minutes. PAV was used to observe the

effect of in-service aging of bitumen i.e. long term aging. The binder was exposed to high temperature of 100°C and a

pressure of 2.1 MPa for 20 hours in pressure aging vessel.

Figure 2. Pressure Aging Vessel (PAV) Figure 3. Rolling Thin Film Oven (RTFO)

2.4. Conventional and Rheological Testing

Penetration, ductility and softening point test were performed on neat and modified binders before and after aging.

To amount the consistency of bitumen at room temperature, penetration test was performed according to AASHTO

specification T49. Softening point of all test samples was determined using Ring and Ball apparatus according to

ASTM D36 and ductility test was performed according to ASTM D 113-17.

Rheological characteristics of neat and modified binders were determined using Dynamic Shear Rheometer (DSR).

Dynamic shear rheometer measures the bitumen properties at different service temperatures, mostly from intermediate

to high. The output of DSR test is in the form of the complex shear modulus (G*) and the phase angle (δ) of bitumen.

DSR is also used to calculate the performance grade of bitumen. DSR used in this research was made of Anton Paar

model 101.25 mm and 8 mm DSR plate geometries were used for this study. Gap between two plates was kept 1 mm

and 2 mm for 25 mm and 8 mm sample respectively. Test was performed at temperature ranges of 45, 52, 58, 64, 70,

76, and 82 and at a constant frequency of 1.59 Hz.

Figure 4. Dynamic Shear Rheometer (DSR)

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1021

2.5. Chemical and Morphological Analysis

Fourier transformed infrared spectroscopy was used to examine the chemical and structural modification of

different samples and to evaluate the influence of aging on modified bitumen. IR radiations were passed through the

sample. The wavelength ranges from 4000 to 400 cm-1

. The resulting spectrum represents the molecular transmission

and absorption of the sample. Morphological analysis of the modified binders before and after aging was carried out

by Scanning Electron Microscope. Samples for the SEM were prepared by putting a drop of PE modified and S

modified asphalt binder on glass slide and then spreading that bitumen uniformly on the surface of the slide in the

form of thin layer to study the dispersion of the modifiers in asphalt binder. The samples were first coated with a thin

gold palladium film and after sputtering, SEM images were taken at different magnifications.

Figure 5. Fourier Transformed Infrared Spectroscopy Figure 6. Scanning Electron Microscope

3. Results and Discussion

In this research, different physical and rheological aging index were used to observe the impact of aging on various

characteristics of neat and modified bitumen. Aging index can be described as the ratio of given property of aged

asphalt binder to the same property of unaged bitumen. Aging index used in this research are Penetration Aging Ratio

(PAR), Softening Point Increment (SPI), Ductility Retained Ratio (DRR), Phase Angle Aging Index (PAAI) and

Complex Modulus Aging Index (CMAI) which can be calculated by the given formulas:

Penetration Aging Ration (PAR) =Aged Penetration Value

Unaged Penetration Value× 100 (2)

Softening Point Increment (SPI) = Aged Softening Point − Unaged Softening Point (3)

Ductility Retained Ration (DRR) =Aged Ductility Value

Unaged Ductility Value× 100 (4)

Phase Angle Aging Index (PAAI) =Aged Phase Angle

Unaged Phase Angle (5)

Complex Modulus Aging Index (CMAI) =Aged Complex Modulus

Unaged Complex Modulus (6)

3.1. Penetration Test Results

Penetration value represents the stiffness and hardening of asphalt binder at normal temperature. Lower the value

of penetration higher the stiffness of bitumen. From Figure 7(a), it can be perceived that with the rise in polyethylene

content from 1 to 5%, the penetration value of bitumen decreased which is an indicative of decrease in fluency and

increase in the consistency of asphalt binder at normal temperature. By addition of 2 to 5% sulphur into the base

binder, the penetration value increases significantly as shown in Figure 7(b) [20]. It indicates that sulphur has a

plasticizing effect on asphalt binder and it has more resistance against thermal cracking especially at low temperatures

[21].

The penetration results after the short term aging and long term aging of the modified bitumen are presented in the

form of PAR in Figures 8(a) and 8(b). PAR of the sulphur modified binder and polyethylene modified binder as shown

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1022

in figures reduced significantly which indicate that the modified bitumen is more resistant to oxidative aging than the

virgin bitumen.

(a) (b)

Figure 7. Penetration value of (a) Polyethylene modified bitumen and (b) Sulphur modified bitumen

(a) (b)

Figure 8. PAR graph of (a) Polyethylene modified bitumen and (b) Sulphur modified bitumen

3.2. Softening Point Test Result

By adding 2% to 5% of polyethylene into the base binder, softening point of the binder increased. 2% addition of

PE resulted in 3% increase in softening point. Even after the aging phase, with the increase in the modifier content, the

high temperature stability of the binder is improved constantly. The impact of aging on the neat and modified binder

can be seen in the form of softening point increment in Figures 9(a) and 9(b). It is generally concluded that the

addition of PE improved the high temperature flowing properties of bitumen and made it more stable against flowing.

It means that the PE modified bitumen has a better high temperature rutting resistance.

For unaged condition, sulphur modified asphalt samples showed a consistent decrease in the softening point which

back the previous findings made on the basis of penetration result that asphalt binder becomes softer on the addition of

sulphur. But after the aging, the hardening level of modified bitumen binder kept on increasing. This can be observed

by the increase in softening temperature after the two aging. Higher softening point asphalt cement is mostly preferred

in hot regions.

0

10

20

30

40

50

60

70

0 2 3 5

Pen

etra

tio

n

Additive rate %

Unaged RTFO aged PAV aged

0

20

40

60

80

100

120

0% 2% 3% 5%

Pen

etra

tio

n

Additive Rate %

Unaged RTFO aged PAV aged

0

10

20

30

40

50

60

70

80

90

0% 2% 3% 5%

PA

R

Additive Rate %

RTFO Aged

PAV Aged

0

10

20

30

40

50

60

70

80

90

0% 2% 3% 5%

PA

R

Additive Rate %

RTFO Aged

PAV Aged

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1023

(a) (b)

Figure 9. Softening Point of (a) Polyethylene modified bitumen and (b) Sulphur modified bitumen

3.3. Penetration Index

It is a quantitative measure of the response of bitumen to the variations in temperature as described by Pfeiffer and

V Doormaal [22]. The type of binder can be identified by its penetration index. Its value generally lies between -3 and

+7. (-3) for highly temperature prone bitumen and (+7) for less temperature prone or highly blown bitumen. Higher PI

values indicate higher temperature resistance. [19]. Generally for road construction, asphalt binder has PI between -2

to +2. Penetration Index (PI) is calculated by following Equation 7 [22].

PI =1952 − 500log (pen) − 200SP

50log (pen) − SP − 120 (7)

Where;

Pen = Penetration

SP = Softening Point

All the PI values obtained in our case are within the normal range of -2 to +2 for road paving application and

shown in Figure 10. The upsurge in PI values indicated the lower temperature vulnerability of modified bitumen.

Figure 10. Comparison of Penetration Index of Polyethylene and Sulphur

3.4. Ductility Test Results

Ductility retained ratio (DRR) is another way of evaluating the impact of aging on the ductile characteristics of

0

10

20

30

40

50

60

0% 2% 3% 5%

Soft

enin

g P

oin

t

Additive Rate %

Unaged RTFO aged PAV aged

0

10

20

30

40

50

60

70

0% 2% 3% 5%

Soft

enin

g P

oin

t

Additive Rate %

Unaged RTFO aged PAV aged

-2

-1.6

-1.2

0% 2% 3% 5%

Pen

etra

tio

n I

nd

ex

Additive rate %

Polyethylene

Sulphur

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1024

bitumen. DRR of polyethylene and Sulphur modified binder is displayed in Figures 11(a) and 11(b) respectively. DRR

in both cases is increasing which represented that the addition of PE and S can reduce deterioration in ductility of

asphalt during aging. At 3% addition of PE the ductility value reduced from 100 to 83 causing a decrease of 25% with

respect to base binder.

(a) (b)

Figure 11. Ductility Retained Ratio for (a) Polyethylene modified bitumen and (b) Sulphur modified bitumen

3.5. FTIR Results

Figures 12(a) and 12(b) demonstrate the IR spectra of neat asphalt binder and binder modified with Polyethylene

and Sulphur. While looking at the spectra of neat asphalt binder we observe the peak in the region of 3200 to 3600

which indicates that OH stretching of alcohol group. Peaks at 2917 and 2846 corresponds to the C-H aliphatic

stretching of alkanes while peak at 1617 refers to C=C aromatic stretching. Peaks in the region of 1720-1750

represents the C=O carbonyl functional group. Peak at 1453 and 1102 corresponds to C-H bending of alkane group

and C-O stretching of secondary alcohol group respectively. The region between 1070-1030 represents the strong S=O

stretching of sulfoxide group. Peaks at 887 and 817 refer to C=C bending of alkene group and peak at 773 indicates C-

S or C-H bending. When we look at the IR spectra of Polyethylene and sulphur modified asphalt binder, very little or

no considerable change of peaks is observed. The intensity of peaks may vary but the range of functional groups more

or less remains the same. It confirms that the modification of asphalt binder with polyethylene and sulphur is purely

physical in nature as it does not change its chemical composition.

(a) Neat and PE modified bitumen (b) Neat and S modified bitumen

Figure 12. IR Spectra of Neat and modified bitumen

IR spectra of neat and modified bitumen after short term aging are shown in Figure 13. It is generalized from the

previous literature view that S=O and C=O are the two functional groups that are responsible for asphalt binder

hardening [23]. Structural index was computed by numerical integrating peak of target functional group and then

dividing it by entire area between 600 cm-1

to 2000 cm-1

[24]. Carbonyl index and sulfoxide index can be calculated by

70

75

80

85

90

95

100

0% 2% 3% 5%

DR

R

Additive Rate %

RTFO Aged

PAV Aged

60

65

70

75

80

85

90

2% 3% 5%

DR

R

Additive Rate %

RTFO AGED

PAV AGED

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1025

numerically integrating the peak of carbonyl and sulfoxide functional groups at 1750 and 1030 respectively and

dividing it by the sum of specific areas as given in Equations 8 and 9.

𝐶𝑎𝑟𝑏𝑜𝑛𝑦𝑙 𝐼𝑛𝑑𝑒𝑥 =𝐴1750

∑ 𝐴 (8)

𝑆𝑢𝑙𝑓𝑜𝑥𝑖𝑑𝑒 𝐼𝑛𝑑𝑒𝑥 =𝐴1030

∑ 𝐴 (9)

Carbonyl and sulfoxide index of the modified binders decreased with respect to neat binder as presented in Figure

14. Carbonyl index decreased by 17% in case of 3% PE modification while it decreased by 23% by addition of 2% S.

Similarly, sulfoxide index decreased by 33% and 42% in case of PE and S respectively. This decrease in the structural

index indicates the greater capability of the modifiers to resist oxidation in asphalt.

(a) Neat and PE modified bitumen (b) Neat and S modified bitumen

Figure 13. IR spectra of bitumen after aging

Figure 14. Structural index of neat and modified binders after RTFO

3.6. SEM Results

Scanning electron microscope test was performed to determine the compatibility and homogeneous dispersion of

modifiers in base bitumen. SEM images of modified binders are shown in Figure 15. Asphalt binder itself is in black

color while white patches shows the presence of modifiers i.e. PE and S. As compared with neat bitumen the

roughness of surface of sulphar modified bitumen was more and it is increased as the percentage of sulphur in bitumen

was increased from 2 to 5%. Tiny particles of sulphur either swollen or dispersed inside the base bitumen can be seen

0

0.01

0.02

0.03

0.04

0.05

0.06

Virgin 60/70 Virgin + 3% PE Virgin + 2% S

Str

uct

ura

l In

dex

Bitumen

Carbonyl Index

Sulfoxide Index

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1026

in SEM images of 2, 3 and 5% sulphur by weight of bitumen. These images also show that bitumen is homogenized

thus it represents the compatibility of modifier with bitumen.

SEM images before and after aging shows good compatibility of modifiers with base binder. Even after the aging,

no phase separation was observed and the surface area was smooth with small irregularities. Therefore it can be

concluded that the modifiers are compatible with the base bitumen and their dispersion is homogenous.

(a) (b)

(c) (d)

(e) (f)

(g) (h)

Figure 15. SEM images (a) 2% S unaged, (b) 2% S aged, (c) 3% PE unaged, (d) 3% PE aged, (e) 3% S, (f) 5% S, (g) 3%

PE, (h) 5% PE

Civil Engineering Journal Vol. 6, No. 5, May, 2020

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3.7. Dynamic Shear Rheometer Analysis

DSR test was performed to check the rheological characteristics of bitumen at intermediate to high temperatures of

45, 52, 58, 64, 70, 76, and 82. Original, modified aged and unaged sample were tested. Complex modulus G* and

phase angle δ were obtained as a result to assess the viscoelastic behavior of different asphalt samples.

3.8. Behavior of Original Bitumen

Complex shear modulus and phase angle represents the behavior of the original bitumen on increasing temperature

range. The G* value of binder decreased with the rise in temperature which shows increase in binder’s stiffness, while

phase angle increased which shows raise in asphalt viscous portion over elastic portion.

3.9. Effect of Modifiers

The effect of modification on the bitumen was obvious as shown in Figure 16. Binder stiffness was increased and

phase angle was decreased by adding polyethylene and sulphur into base binder. It means addition of polyethylene and

sulphur into asphalt resulted in improved elastic behavior of asphalt.

(a) (b)

Figure 16. Variation of (a) Complex Modulus and (b) Phase angle of Neat and Modified bitumen before aging

3.10. Intermediate and High Temperature Performance Characteristics

Impact of aging on the rheological performance of neat and modified bitumen was evaluated in terms of

rheological aging indices. Complex modulus aging index (CMAI) from Equation 10 and phase angle aging index

(PAAI) from Equation 11 were used to assess the aging properties of asphalt binder.

Phase Angle Aging Index (PAAI) =Aged Phase Angle

Unaged Phase Angle (10)

Complex Modulus Aging Index (CMAI) =Aged Complex Modulus

Unaged Complex Modulus (11)

3.11. Complex Modulus Aging Index (CMAI)

The results of CMAI of neat and modified bitumen are presented in Figure 17(a). G* value indicate the bitumen’s

total resistance to permanent deformation. Generally lower value of CMAI indicate higher resistance to aging [25].

Figure 17(a) shows that the CMAI of both modified bitumen is less than the CMAI of neat binder which means that

the PE modified and S modified binder offer higher resistance to oxidative aging, hence improving the rutting

potential of asphalt.

3.12. Phase Angle Aging Index (PAAI)

Phase angle aging index was used to understand the effect of temperature on the behavior of phase angle. Phase

angle present the viscous behavior of asphalt. As the resistance of asphalt against low temperature cracking increases

0

10000

20000

30000

40000

50000

60000

70000

4 6 5 2 5 8 6 4 7 0

Co

mp

lex M

od

ulu

d G

*

Temperature

NEAT

3% PE

2% S

72

74

76

78

80

82

84

86

88

90

46 52 58 64 70 76

Ph

ase

An

gle

Temperature

NEAT

2%S

3% PE

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1028

the value of phase angle also increases [26]. It can be seen in the Figure 17 (b) that at all temperatures, the phase angle

aging index for PE and S modified binder is always greater than the neat binder. It represents that, after aging the

modified asphalt had indicate improved viscous behavior and improved low temperature cracking resistance.

(a) Complex modulus aging index (b) Phase angle aging index

Figure 17. Rheological aging index of neat and modified bitumen

3.13. Determination of Failure Temperature

To determine the failure temperature of neat and modified bitumen, temperature sweep test was performed. SHRP

rutting factor parameter was considered as the failure criteria when it gets below 1 kPa for unaged bitumen and 2.2

kPa for RTFO aged bitumen samples. Failing temperature of the unaged binder improved from 64ºC to 70ºC and

71.5ºC for PE modified and S modified binder respectively. While after the RTFO, the failure temperature increased

to 73.5ºC for polyethylene modified binder and 75ºC for sulphur modified binder. This improvement in the failure

temperature, as shown in Figure 18, indicate the improved rutting resistance of modified binders. Performance grade

of the bitumen increased from PG64 to PG70 in case of polyethylene and sulphur.

Figure 18. Failure Temperature of Neat and Modified Binders

1

1.5

2

2.5

3

3.5

4

4.5

46 52 58 64 70 76

CM

AI

Temperature

NEAT

PE

S

0.8

0.85

0.9

0.95

1

46 52 58 64 70 76P

AA

I

Temperature

NEAT

S

PE

64

70

71.5

67

72.5 73

58

60

62

64

66

68

70

72

74

Virgin 60/70 Virgin +3% PE Virgin + 2% S

Tem

per

atu

re 0

C

Bitumen

Unaged

Rtfo aged

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1029

4. Conclusions

Present study was conducted to determine Polyethylene and Sulphur effect on the aging properties of PARCO

60/70 grade bitumen. Increase in the performance related properties of bitumen increases the service life of roads. To

observe the effect of modifiers on the Performance Grade of bitumen, performance testing of modified and unmodified

bitumen was carried out. The results gave promising benefits of using polyethylene and Sulphur as bitumen modifiers.

Physical properties of the modified bitumen improved as well as the rheological characteristics of modified bitumen in

terms of aging enhanced greatly, thus decreasing the amount of distresses and reducing the pavement cost. The key

findings are concluded as:

• All the physical characteristics of the modified bitumen improved before and after aging;

• Penetration index increased by 10% for Polyethylene modified binder and 36% for Sulphur modification which

indicate better resistance against thermal cracking of the pavement at low temperatures, and lower permanent

(plastic) deformation at high temperatures;

• Addition of Polyethylene and Sulphur into base binder is a Physical process;

• Aging resistance of the modified binder improved which is indicated by decrease in Carbonyl and Sulfoxide

index;

• SEM analysis shows the compatibility of modifiers and their homogenous dispersion in base binder;

• Complex modulus aging index of the modified binders is less than the base binder at all temperatures which

indicate higher resistance against permanent deformation;

• Phase angle aging index of the modified binders is less than the base binder which indicates improved viscous

behavior and improved low temperature cracking resistance;

• Rutting resistance of the modified binders enhanced which is indicated by increase in failure temperatures;

• Performance grade of the binder improved from PG-64 to PG-70.

5. Funding

This research was funded by National University of Science and Technology (NUST), Islamabad.

6. Conflicts of Interest

The authors declare no conflict of interest.

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Available online at www.CivileJournal.org

Civil Engineering Journal

Vol. 6, No. 5, May, 2020

1031

Mechanical Properties of Cement Mortar after Dry–Wet Cycles

and High Temperature

Xiong Liang-Xiao a, b

, Song Xiao-Gang c*

a School of Civil Engineering and Architecture, East China Jiaotong University, Nanchang 330013, P. R. China.

b Guangxi Key Laboratory of Disaster Prevention and Engineering Safety, Guangxi University, Nanning 530004, P. R. China.

c Faculty of Civil and Environmental Engineering, Ningbo University, Ningbo 315211, P. R. China.

Received 01 December 2019; Accepted 19 March 2020

Abstract

The dry–wet cycle and high temperature exposure are important factors affecting the normal use and durability of

concrete structures. The objective of this work is to investigate the mechanical properties of cement mortar specimens

after combinations of dry–wet cycles and high temperature exposures, uniaxial compressive tests on cement mortar

specimens were carried out under the following two sets of conditions: (1) high temperature treatment followed by a dry–

wet cycle and (2) a dry–wet cycle followed by high temperature treatment. The results show that the compressive

strength of specimens increases with the number of dry–wet cycles. After a dry–wet cycle and then a high temperature

treatment procedure, the compressive strength of a specimen will first decrease and then increase with the number of

dry–wet cycles. The strain at the peak stress of cement mortar decreases as the number of dry–wet cycles increases. At

present, there are few research results about the mechanical properties of concrete first after combinations of dry–wet

cycles and high temperature exposures. The work in this paper can enrich the results in this area.

Keywords: Cement Mortar; Mechanical Properties; Uniaxial Compression Test; Dry–Wet Cycle; High Temperature.

1. Introduction

The dry–wet cycle and high temperature exposure are important factors affecting the normal use and durability of

concrete structures. Hydraulic structures, such as piers, in splash zone environments will be affected by the dry–wet

cycle caused by water level changes, and the concrete structure comprising such a pier may also be exposed to fire or

high temperatures.

Many scholars have studied the effects of dry–wet cycles on concrete materials. Li and Shen [1] studied the

deterioration processes of aeolian sand powder concrete under freeze-thaw and dry-wet conditions. Wei et al. [2]

investigated the effect of chloride wet/dry exposure on bonding behavior of BFRP-strengthened concrete beams. Yan

et al. [3] investigated the characteristics of unconfined compression strength and pore distribution of lime-flyash loess

by means of a series of experiments under freeze-thaw cycles or dry-wet cycles. Li et al. [4] analyzed the physical and

mechanical properties of MKPC under dry-wet cycles in 5 wt% Na2SO4 solution. Chen et al. [5] presents an

experimental study on the damage progress of concrete subject to combined sulfate-chloride attack under drying-

wetting cycles and flexural loading. Ma et al. [6] investigated the properties of concrete, including ordinary Portland

concrete and high-performance concrete (HPC), subjected to dry-wet cycles in a variety of salt lake brines. Liu et al.

* Corresponding author: [email protected]

http://dx.doi.org/10.28991/cej-2020-03091526

© 2020 by the authors. Licensee C.E.J, Tehran, Iran. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC-BY) license (http://creativecommons.org/licenses/by/4.0/).

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1032

[7] investigated the coupled effects of external multi-ions and wet-dry cycles in sea water on the evolution of

autogenous selfhealing in cement paste. Yin et al. [8] studied the mechanical properties of textile reinforced concrete

(TRC) under chloride wet-dry and freeze-thaw cycles. Sahmaran et al. [9] found that a dry–wet cycle would accelerate

the degradation rate of the compressive strength of ordinary Portland cement specimens, in sulfate solutions. Gong et

al. [10] firstly conducted experiments on the creep of concrete subjected to dry-wet cycle and sulfate attack. Wu et al.

[11] investigated the transport of chloride ions in concrete under loads and drying- wetting cycles.

A series of related studies have been carried out on the influence of high temperatures on concrete materials. Du et

al. [12] investigated the infrared thermal image inspection, coefficient of thermal conductivity, apparent density, and

compressive strength test on C80 high-strength concrete (HSC) in the presence and absence of polypropylene fibers

under completely heated conditions. Khan and Abbas [13] presented the behavior of high volume fly ash concrete at

varying peak temperatures. Meng et al. [14] studied the triaxial compressive properties of recycled aggregate concrete

(RAC) after high temperature. Li et al. [15] investigated the static and dynamic mechanical properties of concrete

before and after high temperature exposure. Liu et al. [16] experimentally and analytically investigated the residual

strength of SRC cross-shaped columns after exposure to high temperatures. Zhai et al. [17] conducted intensive SHPB

tests and corresponding quasi-static tests were to study the strain-rate effects on the normal weight concrete after high

temperature. Arioz [18] found that as the high temperature exposure was increased, the quality and the relative

strength of the concrete specimens decreased significantly. Ma et al. [19] discussed the effects of their tested high

temperature methods on the mechanical properties of concrete after exposure, through several sets of experiments.

Although there have been many studies on the mechanical properties of concrete after dry–wet cycles and after

high temperature exposures, most of the results only consider single variables; experimental results for pairing dry–

wet cycles and high temperatures are rare. The requirements of environments where actual concrete structures are used

are complex and variable, ranging from the concrete structures being subject to fire after long periods of continuous

wet, due to rainy seasons; although some concrete structures are exposed to fire, they often stay in use due to the

damage not being obvious. Such structures could be subjected to dry–wet cycles caused by adverse weather in their

subsequent service phases. Therefore, the mechanical properties of concrete after the combination of dry–wet cycles

and high temperatures must be studied, that is, the mechanical characteristics of concrete first after high temperatures

followed by dry–wet cycles and vice versa.

In order to study the mechanical properties of the cement mortar after these combinations of dry–wet cycles and

high temperature exposures, in depth, the cement mortar specimens produced from one batch have been subjected to

the dry–wet cycle first and then to the high temperature treatment and vice versa. Uniaxial compression tests were

carried out on the specimens, and their compressive strength, peak strain, and elastic moduli were investigated in light

of the number of dry–wet cycles and high temperature exposures they had experienced.

2. Research Methodology

The program of this work can be seen in by a flow chart as shown in Figure 1.

No. 425 Portland cement

Compressive strength

Cubic cement mortar specimens, dimension of

70.7×70.7×70.7 mm

Experimental Works

ISO standard sand

Cement/sand ratio is 1:2

Water/cement ratio is 0.65

Uniaxial compression test

Elastic modulus

Conclusions

First after high temperature treatment

followed by dry-wet-cycle

First after dry-wet-cycle followed

by high temperature treatment

Strain at the peak stress

Figure 1. Flow Chart for Research Methodology

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1033

3. Test Setup

3.1. Test Instruments

The uniaxial compression test was conducted using the WAW600 universal testing machine. The machine had a

maximum displacement loading rate of 60 mm/min and a maximum load of 600 kN. The oven used was the type 101–

2 electric heating, constant temperature, blast drying box produced by Shangyu Geotechnical Instrument Co., Ltd. The

high temperature box used was the GWM-1100 electric heating, high temperature test box produced by Changchun

Fangrui Technology Co., Ltd.

3.2. Sample Preparation

The cement was 42.5 grade ordinary Portland cement. The fine aggregate was freshwater river sand produced in

Ningbo; it belongs to the medium sand category, with particle-level matching grid. The mixing water used was

ordinary tap water.

The size of each specimen was 70.7×70.7×70.7 mm, and the water/cement ratio was 0.55. A total of 138 specimens

were prepared, and their molds were removed after being placed at room temperature for 24 h. All cement mortar

specimens were subjected to the next step of the research after 28 days of standard curing.

3.3. Experiment Procedure

(1) High temperature treatment method: the specimens were placed in a high temperature box, after which the

temperature was raised to a preset temperature (200, 300, 400) at a heating rate of 10°C/min. The temperature

was then kept stable for 2 h, when the high temperature box was turned off to stop heating. The specimens were

removed from the box after naturally cooling for 24 h.

(2) Dry–wet circulation treatment method: the specimens were immersed in water for 15 h, after which the

specimens were taken out and any surface moisture dried with a dry towel. They were then left to stand for 30 min

before being transferred to the 40°C oven for 8 h. They were then removed and allowed to cool to room temperature

for 30 min, soaked in water. The cycle was then repeated with each cycle taking 24 h.

(3) Uniaxial compression test: All the specimens were allowed to stand at room temperature for 7 days after their

high temperature and dry–wet cycle treatments. They were then subjected to uniaxial compression testing. The test

loading mode adopted displacement control with a loading rate of 0.4242 mm/min.

The numbers of the three specimens without high temperature and dry–wet circulation treatments are set as # 0.65-

RM-1, # 0.65-RM-2 and # 0.65-RM-3. The loading-time curves of these three specimens are shown in Figure 2.

Figure 2. The loading-time curves of specimens

3.4. Test Group

A total of 18 specimens underwent dry–wet cycles at room temperature. A total of 54 specimens had the high

temperature exposure first and then the dry–wet cycle. A total of 54 specimens had the dry–wet cycle before the high

temperature phase. A total of 12 specimens underwent only the high temperature treatment at different temperatures.

Three specimens were used for each test, and the compressive strength, peak strain, and elastic moduli values used

here are the average values for these three specimens.

0 50 100 150 200

0

50

100

150

200

Fo

rce

/ k

N

Time / s

# 0.65-RM-1

# 0.65-RM-2

# 0.65-RM-3

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1034

4. Test Results and Analysis

4.1. Change Law of Compressive Strength

The change law of compressive strength of the specimens with the number of dry–wet cycles is shown in Figure 3.

The compressive strengths of specimens after high temperature and then the dry–wet cycle are shown in Figure 3(a),

and the compressive strengths of specimens in the dry–wet cycle and then high temperature group are shown in Figure

3(b).

(a) First after high temperature and then the dry–wet cycle (b) First after dry–wet cycle and then the high temperature

Figure 3. Variation of compressive strength of specimen with dry–wet cycling times

According to Figure 3 (a):

When the number of dry–wet cycles is nil, the specimens having the high temperature and then the dry–wet cycle

only show the effects of the high temperature. The compressive strength of specimens gradually decreases as the

temperature rises from room temperature (25°C) to the highest temperature used, which is 400°C.

The compressive strength of the specimens that underwent the dry–wet cycles at room temperature increased with

the number of dry–wet cycles, mainly because the clinker mineral hydrates in the cement increase with the length of

the time spent in sufficient water. The gel is filled with capillary pores making the inside of the specimen more

compact, thereby improving the compressive strength of the specimen.

After exposure to high temperatures of 200 and 400°C, the compressive strength of the specimen first decreases

and then increases slightly, as the number of dry–wet cycles increases. However, the compressive strength remains

lower than that of a specimen having the same number of dry–wet cycles at room temperature.

Due to the different thermal expansion coefficients of the cement slurry and the medium sand, microcracks will

appear on the surface of the specimen after experiencing high temperatures. The repeated action of the dry–wet cycle

will increase these microcracks. When the number of cycles is small, the continuous increase of microcracks will

reduce the overall structural strength of the specimen; as the number of cycles increases, the water gradually enters the

inside of the specimen along the numerous microcracks. The water reacts with the unhydrated cement particles. The

chemical action brings a certain degree of improvement to the compressive strength, but the high temperatures have a

degrading effect on the compressive strength of the specimen, so the compressive strength is still lower than that of a

specimen that had the same number of dry–wet cycles at room temperature.

After the same number of dry–wet cycles, the compressive strength of a specimen that underwent dry–wet cycles at

room temperature is greater than the compressive strength of a specimen that underwent high temperature exposure

and then dry–wet cycle process, with the same number of cycles. At the same time, the compressive strengths of the

specimens experiencing high temperatures of 300°C and 400°C, before their dry–wet cycles, are greater than the

compressive strength of a specimen experiencing a high temperature of 200°C before the same number of dry–wet

cycles. This is because when a specimen is subjected to high temperatures of 300°C and 400°C, the number of

microcracks and pores that are formed inside the specimen is greater than that of a specimen subjected to a high

temperature of 200°C. The dry–wet cycle allows water to enter the specimen through the pores and cracks, so that the

degree of hydration of the cement clinker mineral is increased, increasing the compressive strength too. The denser the

distribution of pores inside the specimen, the more the degree of hydration increases, and the greater the strength

increase. However, the high temperature still does certain damage, deteriorating the specimen, so the compressive

strength is still lower than that of specimen that has not been subjected to the intense heating.

Civil Engineering Journal Vol. 6, No. 5, May, 2020

1035

According to Figure 3 (b):

When specimens have their dry–wet cycles before experiencing high temperatures of 200°C and 300°C, their

compressive strengths generally first decrease slightly and then increase with the number of dry–wet cycles. There are

two reasons: (1) When the number of dry–wet cycles is relatively small, the dry–wet cycling causes microcracks in the

cement mortar specimen, reducing the compressive strength of the specimen. (2) When the number of dry–wet cycles

is larger, as the number of cycles continues to increase and microcracks continue to expand inside the specimen, the

moisture that is present reacts with the unhydrated cement particles, continuously filling the pores, improving the

compressive strength of the specimen.

The comparisons of the compressive strengths of the specimens that experienced the dry–wet cycles first with the

strengths of those that experienced the high temperatures first, for the same high temperatures, are shown in Figure 4.

(a) 200 (b) 300

(c) 400

Figure 4. Comparisons of the compressive strengths of the specimens that experienced the dry–wet cycles first with the

strengths of those that experienced the high temperatures first, for the same high temperatures

In general, when the number of cycles is the same, the compressive strengths of specimens that experienced the

dry–wet cycles first are higher than the strengths of those that experienced the high temperatures first, for the same

high temperature.

4.2. Variation Law of Strain at the Peak Stress

The change law of strain at the peak stress of the specimens, with the number of dry–wet cycles, is shown in Figure

5. The strains at the peak stress of the specimens that experienced high temperatures before their dry–wet cycles are

shown in Figure 5(a), and the strains at the peak stress of the specimens that had their dry–wet cycles before

experiencing high temperatures are shown in Figure 5(b).

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(a) First after high temperature and then the dry–wet cycle (b) First after dry–wet cycle and then the high temperature

Figure 5. Variation of the strain at peak stress of specimens with dry–wet cycle times

When the specimen only experienced high temperature, when the number of dry–wet cycles was nil, the strain at

the peak stress of the specimen increases with the increase in temperature; the strain at the peak stress after the high

temperature of 400°C is double that of the room temperature specimen, indicating that the high temperature greatly

improves the ductility of the specimen. The strains at the peak stress of specimens that experienced the dry–wet cycles

before heating and those that experienced the high temperatures before the dry–wet cycles both showed decreasing

trends with the increase in the number of cycles, indicating that the dry–wet cycling reduces the ductility of the

specimens.

4.3. Variation Law of Modulus of Elastic Modulus

In this paper, the secant elastic modulus was calculated at a peak stress of 30%, and the corresponding strain was

taken as the elastic modulus of the specimen. The change law of the elastic moduli of specimens with the number of

dry–wet cycles are shown in Figure 6. The elastic moduli of specimens that experienced high temperatures before their

dry–wet cycles are shown in Figure 6 (a), and the elastic moduli of specimens that had their dry–wet cycles first are

shown in Figure 6 (b).

(a) First after high temperature and then the dry–wet cycle (b) First after dry–wet cycle and then the high temperature

Figure 6. Variation of specimen elastic modulus with dry–wet cycling times

According to Figure 6(a):

When the number of dry–wet cycles was nil, as the high temperature gradually increased from room temperature

(25°C) to 400°C, the elastic modulus of the specimens gradually decreased. This is consistent with the findings of

Sahmaran et al. [9]. The elastic moduli of the specimens that experienced high temperatures of 200°C, before their

dry–wet cycles, first decreased and then increased as the number of dry–wet cycles increased. The elastic moduli of

the specimens that experienced high temperatures of 300°C and 400°C, before their dry–wet cycles, increased with the

number of dry–wet cycles.

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According to Figure 6(b):

The elastic moduli of specimens after dry–wet cycles at room temperature, and dry–wet cycles before experiencing

a high temperature of 200°C, generally first decreased and then increased as the number of dry–wet cycles increased.

For the high temperatures of 300°C and 400°C before the dry–wet cycles, the elastic moduli of specimens increased

with the number of dry–wet cycles.

5. Conclusions

When the specimens only experienced high temperatures, as the temperature increased gradually from room

temperature (25°C) to 400°C, the compressive strengths and elastic moduli of the specimens gradually decreased,

while the strain at the peak stress gradually increased. The compressive strengths of the specimens that underwent

dry–wet cycling at room temperature increased with the increase in the number of dry–wet cycles. The

compressive strength of the specimens that experienced both heating and dry–wet cycles, in either order,

generally first decreased then increased as the number of dry–wet cycles increased.

The peak strains of the heating before dry and wet cycles group and the dry and wet cycles before heating group

decreased with the increase in the number of cycles.

The elastic moduli of specimens that underwent dry–wet cycling at room temperature, and those that had dry–wet

cycles followed by a high temperature of 200°C, generally first decreased and then increased as the number of

dry–wet cycles increased. The elastic moduli of specimens that had dry–wet cycles followed by high temperatures

of 300°C and 400°C increased with the number of dry–wet cycles.

6. Funding Acknowledgements

This work was supported by the Systematic Project of Guangxi Key Laboratory of Disaster Prevention and

Engineering Safety (Grant No. 2019ZDK051), and the Open Research Fund of State Key Laboratory of Simulation

and Regulation of Water Cycle in River Basin(China Institute of Water Resources and Hydropower Research, (Grant

No. IWHR-SKL-201708).

7. Conflicts of Interest

The authors declare no conflict of interest.

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