75504
2712-8172
Magazine of Civil Engineering
19
2
142
2026
1-114
RAR
RUS
14201-14201
University of Anbar
Ftaikhan
Ahmed
ahm22e1008@uoanbar.edu.iq
Baghdad, Iraq
56184497100
0000-0001-6180-8837
University of Anbar
Al-Sharrad
Muayad
muayad.alsharrad@uoanbar.edu.iq
Ramadi, Iraq
Effect of compaction pressure and wheat straw inclusion on geopolymer-stabilized rammed earth behavior
This paper presents the effect of compaction pressure level and wheat straw addition on the durability, mechanical behavior, and thermal conductivity of geopolymer stabilized rammed earth. Rammed earth specimens were prepared by static compaction to 5, 10, and 25 MPa of mixtures containing predefined amounts of sand, silt, clay, and wheat straw, stabilized with fly ash geopolymer. A number of unstabilized specimens made of the raw materials were also prepared for comparison. These specimens were cured inside plastic bags at 35 °C so that the least energy consumption is achieved. The durability was investigated by performing a dip test and spray test. The results of the geopolymer stabilized specimens demonstrated an excellent resistance to erosion by water, unlike the unstabilized specimens, which almost failed completely. The mechanical behavior was evaluated by performing unconfined compression test. The results indicated that material’s stiffness and strength increased considerably with increasing compaction pressure and curing age, with the majority of the increase occurring during the first month of curing. Compressive strength values of 4.2 and 10 MPa were recorded from tests on stabilized specimens compacted to 5 and 25 MPa, respectively, then cured for 60 days. These figures are promising, keeping that, a threshold unconfined compressive strength of 1–2 MPa is typically acceptable by many building codes. A relatively low thermal conductivity of about 0.35–0.5 W/(K.m) was recorded from the hot wire method on specimens prepared from various mixtures, suggesting that the stabilized rammed earth outperforms most of the traditional building materials such as concrete. The inclusion of wheat straw improved material’s ductility by increasing strains, at which shear failure occurs. However, this inclusion led to undesirable reduction in strength and stiffness over the first two months of curing and almost no change in thermal conductivity, with respect to those obtained on the stabilized specimens. This response was attributed to material wise incompatibilities.
1.Introduction
Rammed earth is a technique that uses sand, silt, clay, and sometimes gravel as raw materials by compacting them into forms to construct walls. Since ancient times, humans have used earthen materials in construction. They are available at a low cost and have good mechanical and insulation properties so they have been used as an excellent solution for construction throughout history [1–3]. Recently, rammed earth construction has become more popular in the sustainability arena since the materials are still inexpensive and locally available [4], and well suitable for mixing with stabilizers such as lime [5, 6], waste materials, and factory side products such as fly ash [7, 8], or similar mixtures [9–15].
Although rammed earth has limited compressive strength, it has often been used for structural purposes [16]. For this reason, rammed earth is often stabilized using Portland cement. The use of Portland cement has a negative environmental impact, as cement production consumes a lot of energy and water, which represents 9 % of the global consumption of industrial water [17, 18], and it is responsible for producing 5 % of harmful gases every year [19]. Hence, it is indispensable to search for alternative stabilizers such as fly ash, crushed bricks, or natural additives (e.g., natural fibers) [18, 20]. The use of stabilizing materials has a positive effect on improving hygroscopic properties and reducing thermal conductivity [21]. According to [17, 22], thermal conductivity decreases when the wheat straw is added for rammed stabilization. On the other hand, the use of wheat straw reduces compressive strength compared to samples that do not contain wheat straw.
The current study explores the effect of compaction pressure and the addition of wheat straw on several properties of rammed earth, i.e., durability, strength, and thermal conductivity. In practice, reducing the compaction pressure to a level as low as 5 MPa has many advantages, including lower production costs and improved thermal insulation.
2.Methods and Materials
2.1.Materials
Soils. The soils were obtained from quarries near Ramadi, where sufficient quantities of coarse and fine soils were collected and transported inside plastic bags to the soil laboratory at the University of Anbar (Fig. 1a and 1b). The soils were dried at 110 °C, crushed, then samples were taken for classification. Table 1 shows index properties of these soils. Soil 1 was classified as poorly graded sand, while Soil 2 was classified as fat clay.
Figure 1. Materials used in this work: a) Soil 1; b) Soil 2;
c) crushed wheat straw; d) fly ash and NaOH.
Table 1. Index properties of Soil 1 and Soil 2.
Property
Soil 1
Soil 2
Specification
Gravel (>4.75 mm, %)
0
0
ASTM D422-2007 [23]
Sand (4.75–0.075 mm, %)
100
0
Silt (0.075–0.005 mm, %)
0
35
Clay (Compaction Characteristics
(a)
(b)
Figure 5. Compaction characteristics of: a) raw material; b) stabilized material.
< >Durability
Time (min)
of test
Depth of erosion, D (mm)
Criteria
Erodibility index
Unstabilized material
Stabilized material
15
10
≈ 0.0
0 ≤ D < 20
20 ≤ D < 50
50 ≤ D < 90
90 ≤ D < 120
D ≤ 120
1
2
3
4
5 (fail)
30
18
45
24
60
30
< >Unconfined Compression
(a)
(b)
Figure 7. Typical stress-strain curves after 30 days of curing for:
a) stabilized material without wheat straw; b) stabilized material with wheat straw.
Figure 8. Variation of the average compressive strength with curing age for:
a) stabilized material without wheat straw; b) stabilized material with wheat straw.
Figs. 9 and 10 analyze the effect of compaction pressure as well as curing age on the elastic modulus and compressive strength, respectively. The geopolymer acted as a cementitious material, which provided additional bonding to sand particles and, by other words, restricting the relative movement at the interparticle contacts. As a result, the geopolymer-stabilized specimens exhibited very much stiffer response than raw material compacted to the same pressure. For example, the specimen of raw material, which was compacted to 25 MPa, displayed an elastic modulus of only about 16 MPa, compared to 1306 and 560 MPa displayed by the corresponding specimens without and with wheat straw, respectively. This also indicates that wheat straw inclusion led to a large decrease in the elastic modulus. The elastic modulus increased substantially over the first month with curing then showed lesser increase by the end of the second month.
Inspection of Fig. 10 indicates that compressive strength values of the stabilized specimens were remarkably higher than those made of the raw material, owing to the cementation function of the geopolymer. Quantitatively, an increase of about one to two orders of magnitude can always be observed in the compressive strength.
(a)
(b)
Figure 9. Variation of the elastic modulus with compaction pressure:
a) without wheat straw b) with wheat straw.
(a)
(b)
Figure 10. Variation of the compressive strength with compaction pressure:
a) without wheat straw b) with wheat straw.
< >Thermal Conductivity SEM ImagingOverall, the stabilized rammed earth examined in this work demonstrated promising durability, mechanical, and thermal properties. For instance, a threshold unconfined compressive strength of 1–2 MPa is recommended by many building codes. In addition, the average thermal conductivity of the geopolymer stabilized rammed earth is much lower than that of many building materials such as ordinary concrete, which has a thermal conductivity of about 1.5–3.5 W/(m.K).
10.34910/MCE.142.1
624
rammed earth
compressive strength
durability
wheat straw
thermal conductivity
fly ash geopolymer
https://engstroy.spbstu.ru/article/2026.142.1/
RAR
RUS
14202-14202
57366175000
0000-0003-1808-0208
Don State Technical University
Kaklyugin
Alexandr
kaklugin@gmail.com
Rostov-on-Don, Russian Federation
57204555855
0000-0001-8968-2543
Don State Technical University
Kastornykh
Lyubov
likas9@mail.ru
Rostov-on-Don, Russian Federation
0000-0002-9468-3062
Brest State A.S. Pushkin University
Stupen
Nonna
chemskorp@yandex.ru
Brest, the Republic of Belarus
0000-0002-9860-5774
Brest State A.S. Pushkin University
Kovalenko
Viktor
kvv0407@rambler.ru
Brest, the Republic of Belarus
Regulation of structure formation of pressed composites based on the modified gypsum binder
The main disadvantage of gypsum and gypsum concrete products molded by casting, vibrating, rolling, and pressing methods is low water resistance, which manifests itself in a significant decrease of strength at humidification. The authors have shown the possibility of increasing the strength and water resistance of pressed gypsum products by modifying their structure with an additional crystallization framework of sparingly soluble calcium hydrogen phosphate dihydrate. The purpose of the research is to obtain an adjustment equation that makes it possible to determine the rational dosages of modifying additives depending on the values of the specified strength and water resistance of pressed gypsum composites, as well as to assign the optimal duration of mixing the semi-dry molding mixture until the sealing pressure is applied to it. Using mathematical experimental planning methods, there was investigated the effect of the mixing time of the molding mixture and the dosages of the modifying additives on the basic physical and mechanical properties of pressed gypsum composites compacted at a pressure of 40 MPa. Ammonium dihydrogen phosphate and carbonate-containing sludge from the chemical water treatment of thermal power plant were used as modifiers. It has been revealed that in the proposed technology, along with the dosages of modifying additives, the physical and mechanical properties of the material are significantly affected by the duration of mixing the molding mixture, during which chemical interaction occurs between its components, which increases the strength and water resistance of gypsum modified pressed composites. Experimental and statistical models of the most important technical characteristics of the proposed material have been developed, depending on the main prescription and technological factors, which make it possible to determine the conditions for obtaining pressed gypsum products with specified properties. An adjustment equation has been obtained that makes it possible to establish rational dosages of modifying additives and the optimal mixing time of the molding mixture at a given value of the softening coefficient.
1.Introduction
Gypsum construction products molded from semi-dry mixtures under high pressures are characterized by a simple production technology and have a number of valuable properties. First of all, they have a high compressive strength (up to 60–70 MPa), which can be attained even with low-grade gypsum binders, and they do not require artificial drying [1, 2]. However, they are united with other types of gypsum and gypsum concrete products formed by casting, vibrating, or rolling methods by low water resistance, which manifests itself in a significant decrease in strength when wet [3, 4]. This is explained by the dissolution of calcium sulfate dihydrate at the contact points of crystalline junctions in its structure, as well as the wedging effect of aqueous films adsorbing in microcracks and capillary pores during water saturation [5, 6].
The hardened gypsum binder’s structure can be made more resistant to the dissolving and wedging action of water, as many researchers claim, by adding an additional crystallization framework composed of sparingly soluble compounds [7–9].The results of our previous studies show that the water resistance of pressed gypsum composites in-creases significantly in the case of the formation of shielding protective films from sparingly soluble calcium hydrogen phosphate dihydrate on the elements of their fine-crystalline structure due to the modification of gypsum binder by the addition of ammonium dihydrogen phosphate NH4H2PO4 [10, 11]. In the technology of gypsum products made by casting, chemical additives that form sparingly soluble phase films, such as phosphate films, on the surface of grains of hydrated neoplasms are usually used in small dosages to slow down the setting of gypsum binders [12–14]. The pressing method allows for a significant increase in the dosage of the chemical additive because right after molding the material has a high structural strength. Excessive inhibition of the processes of structure formation in this case is not dangerous from the point of view of in-house transportation of freshly molded products. However, in the future, more complete shielding of the formed fine-crystalline structure of the material with films of sparingly soluble compounds will ensure an increase in its water resistance.
In gypsum mixtures with a reduced water content, various fine fillers can have a positive effect on the processes of binder formation. At the same time, in order to increase the strength and water resistance of gypsum products, it is appropriate to use fillers that can form water-resistant complexes with gypsum, as well as with chemical additives used [15–17]. In this case, the water adsorbed by the filler grains can serve as a reserve for the hydration of gypsum binder in a freshly molded product, and calcium sulfate dihydrate, which crystallizes with some increase in volume, will be deposited in the pores of the material, thereby healing defects in its structure [18–20]. At the same time, choosing the appropriate water-binding ratio is essential. It is necessary to ensure, on the one hand, proper compaction of the press powder and the absence of water separation, and on the other hand, complete hydration of the gypsum binder and, consequently, stability of properties and the absence of volumetric deformations of the artificial stone-like material over time. Given the aforementioned, in previous research, we used carbonate-containing sludge from the chemical water treatment of thermal power plants as a filler. The possibility of replacing a significant part of the gypsum binder (up to 60 %) with a carbonate-containing filler was established. Lower product costs and improved environmental conditions around thermal power plants are possible additional outcomes of using a secondary resource as a modifying filler [10, 11]. However, the analysis of the results of the pre-experiments revealed that in our proposed method of modifying the composition and structure of pressed gypsum composites, one of the most important factors is the appropriate selection of the molding mixture preparation time. It is during the mixing of gypsum binder with a modifying filler sealed with an aqueous solution of NH4H2PO4 that the binder starts to hydrate and the chemical additive interacts with both CaSO4·0.5H2O binder and CaCO3 sludge. The ongoing chemical processes lead to the formation of films of sparingly soluble calcium hydrogen phosphate dihydrate CaHPO4·2H2O (brushite) on the surface of semi-aqueous gypsum grains. This compound is isomorphic to dihydrate gypsum, but it is more than ten times less soluble. The formation of calcium hydrogen phosphate dihydrate also occurs on the surface of the grains of the modifying filler due to the interaction of NH4H2PO4 with CaCO3 of the chemical water treatment sludge. Further, these films protect the crystallization contacts of already hardened composites from dissolution at humidification, which largely increases their water resistance.
The chemical and physicochemical processes occurring when the molding mixture is mixed are described in more detail in [21, 22]. However, the nature and intensity of these processes, as well as the energy state of the system, as demonstrated by the research findings of the scientific school of Professor O.P. Mchedlov-Petrosyan [23, 24], must be coordinated with the moment of technological impact on the molding mixture obtained after mixing. In this instance, the use of modifying additives is likely to have the biggest impact. In the technology of molding gypsum products by pressing, such an effect is reached by, first of all, the application of sealing pressure to the prepared mixture. Pressing at the moment of changing the hydration stage of the binder activates and enhances the structural processes occurring in the system, ensures the formation of a more durable and water-resistant crystallization framework at further stages of structure formation.
The purpose of this work is to obtain an adjustment equation that makes it possible to determine the rational dosages of modifying additives depending on the values of the required strength and water resistance of pressed gypsum composites, as well as to assign the optimal mixing time of the semi-dry molding mixture until sealing pressure is applied to it.
To achieve the goal, it is necessary to consistently solve the following tasks:
to study the effect of the duration of mixing the molding mixture on the basic physical and mechanical properties of pressed gypsum composites at various consumption rates of modifying additives;
to establish rational intervals for varying the mixing time of the molding mixture, the dosages of the chemical additive and the modifying filler, and to perform experimental and statistical modeling of the basic physical and mechanical properties of the final product, depending on these factors;
to analyze the experimental statistical models and with these results to obtain an adjustment equation that enables the control of the structure formation processes of pressed gypsum and demonstrates that each of the consumption rates of carbonate filler and ammonium dihydrogen phosphate corresponds to its optimal mixing time.
2.Methods
In experimental studies, we used G-5 B II grade gypsum binder according to Russian State Standard GOST 125-2018 “Gypsum binders. Specifications.” Finely dispersed carbonate-containing sludge from the chemical water treatment of a thermal power plant in combination with a chemical additive, ammonium dihydrogen phosphate, was used as a modifying filler in the composition of molding mixtures.
The carbonate-containing filler used in these studies is waste generated in the technological process of softening the water of the Don River with slaked lime and soda at a thermal power plant. As a result of this preparation, the calcium and magnesium salts present in the river water are converted mainly into carbonates of these metals. The resulting sludge is a secondary resource, which is commonly referred to as sludge from chemical water treatment. After drying, it is a fine powder with a specific surface area of 780 m2/kg. Ammonium dihydrogen phosphate (NH4H2PO4) is a salt formed as a result of the neutralization of orthophosphoric acid by ammonia. Details regarding the additives, their mechanism of action, and effects are provided in [10, 11].
When molding gypsum products by casting or vibration methods, it is impossible to modify the carbonate filler with acid additives directly during the preparation of the molding mixture, since the carbon dioxide released during this process causes the gypsum dough to become porous. This feature is used in the manufacture of gypsum products by casting. In a semi-dry mixture of gypsum binder and finely dispersed calcium carbonate, suitable for the manufacture of products by pressing, carbon dioxide can be freely released into the environment [21, 22]. In this regard, selecting the appropriate molding mixture preparation method and duration based on the ongoing chemical processes is a crucial task [25, 26].
The components of the molding mixture (gypsum binder and sludge sealed with ammonium dihydrogen phosphate solution) were mixed in a laboratory slider mixer. The completeness of neutralization of the acidic solution of the chemical additive during mixing was controlled by changes of the pH of the aqueous extracts (1:5) using a potentiometric method on a universal EV-74 ionomer with platinum and glass electrodes.
Control samples, cylinders with a diameter and height of 50.5 mm, were formed from the press powder obtained after mixing at a pressure of 40 MPa in special molds. The prepared samples were stored for 24 hours in air-dry conditions and then dried in a drying cabinet at a temperature of (55±5) °C to a constant weight. Half of the samples from each batch were immersed in water for 24 hours. The dried and water-saturated samples were tested for compressive strength and, based on the results obtained, the coefficient of softening was calculated as the ratio of the strengths of the material in the dry and water-saturated state.
Experimental and statistical modeling of the technical characteristics of gypsum modified pressed composites was carried out using mathematical experimental planning methods. For these purposes, a complete factorial experiment on the B3 Box composition plan was carried out [27]. Mathematical models in the form of a second-order polynomial that adequately describe the technical characteristics of the proposed material were constructed with the help of our original software products. The adequacy of the obtained models was assessed by the value of the Fisher criterion calculated for them [27].
3.Results and Discussion
As already mentioned in the introduction of this paper, as a result of our earlier experimental findings [10, 11], there was revealed the possibility of increasing the strength and water resistance of pressed gypsum composites by modifying their composition and structure with additives of carbonate sludge from chemical water treatment of thermal power plants and ammonium dihydrogen phosphate. The proposed technology does not require the development and manufacture of special molds and devices. The amount of time the mixture is held under pressure does not exceed the time spent when pressing products based on other mineral binders. It was found that the time of mixing is an important factor in the proposed technology to produce gypsum products, providing for the compaction of a semi-dry molding mixture by pressing. At the same time, an important element of the mechanism of action of a chemical additive is its interaction not only with calcium sulfate hemihydrate of the gypsum binder but also with calcium carbonate of the filler. An aqueous solution of the chemical modifier has an acidic environment (pH ≈ 3.7). This is explained by the fact that when NH4H2PO4 is dissolved in water, along with the hydrolysis of this salt, dissociation of the dihydrogen phosphate ion occurs:
The hydrogen ions present in the solvent water interact with the CaCO3 of the filler, which leads to the formation of calcium hydrogen phosphate dihydrate on the surface of its grains.
Due to the high concentration of calcium carbonate in the mixture, these processes gradually lead to a change in the pH of its medium from acidic to neutral. Control of the neutralization of an aqueous solution of ammonium dihydrogen phosphate during its mixing with the dry components of the molding mixture was carried out according to the procedure described above. The mixture was sampled for pH measurement after every minute of stirring. The pH change of the mixture is shown in Table 1.
Table 1. Changes in the pH of components and mixtures depending on the duration of mixing.
Number
of composition
Content of components, %
Consumption of NH4H2PO4, %
of the mass of the dry components of the mixture
Duration of mixing, min
рН of
water
extracts
gypsum binder
sludge
from chemical water treatment of thermal power plants
1
100
0
0
1 and more
7.15
2
100
0
2
1
2
3 and more
6.80
7.10
7.30
3
0
100
0
1 and more
9.00
4
80
20
0
1
2 and more
8.10
8.30
5
80
20
2
1
2 and more
7.00
7.40
6
60
40
2
1
2 and more
7.20
9.00
As seen from Table 1, after 2 minutes of mixing, the pH of the mixture rises to approximately neutral (pH = 7.00). This indicates a high rate of reactions between ammonium dihydrogen phosphate, calcium sulfate, and calcium carbonate. As a result of chemical processes occurring during mixing of the molding mixture, there appear hydrate neoplasms on the resulting microcrystals and in the places of their contact with the surface of the filler grains – sparingly soluble protective films of calcium hydrogen phosphate dihydrate (CaHPO4·2H2O). These films play an important role in the organization of the fine-crystalline structure of the artificial stone-like material and later complicate the dissolution of crystallization contacts at humidification, thereby providing increased water resistance.
In the proposed method of increasing the strength and water resistance of pressed gypsum composites, it should be taken into account that at the stage of their structure formation, a significant effect is exerted by slowing down the hydration of gypsum binder due to the formation of shielding films of sparingly soluble calcium hydrogen phosphate dihydrate on the surface of its grains [10, 21, 22].
Fig. 1 shows the dependence of the compressive strength of dried samples, and Fig. 2 shows the dependence of water-saturated samples made by pressing from various compositions on the duration of mixing after which they were molded.
The duration of mixing varied from 2 to 20 minutes. All the compositions studied were prepared with a water-solids ratio of W/S = 0.19, which ensures complete hydration of semi-aqueous gypsum, and the samples were molded at a pressing pressure of P = 40 MPa.
As seen from Figs. 1a and 2a, the strength of the pure gypsum binder material is significantly affected by the quantitative ratio between the additives introduced and the duration of mixing the mixture. Longer mixing of the mixture leads to a significant decrease in the strength of dried and, to an even greater extent, water-saturated samples. This is due to the high rate of hydration of the gypsum binder, which is largely completed already during the mixing process. As the mixing time increases, more calcium sulfate dihydrate is formed and the amount of active hemihydrate decreases, the hydration of which in the molded product mainly determines the strength of the resulting material structure.
With the introduction of 1 % NH4H2PO4 additive (Figs. 1a and 2a, Curve 2), the dependence of the strength of the material on the mixing duration is of a different nature. The samples molded after 2 minutes of stirring have approximately the same strength as the samples without the addition of ammonium dihydrogen phosphate. An increase in the duration of mixing leads to an increase in the strength of the material. The samples molded from this composition have the maximum strength when the mixing time of the molding mixture is 12–14 minutes. With a longer mixing duration, the strength of the samples decreases. In compositions that do not contain a carbonate filler, an increase in the consumption of the NH4H2PO4 additive negatively affects the strength and water resistance of the samples made.
а)
b)
c)
Figure 1. The dependence of the compressive strength of dry samples on the duration
of mixing of the molding mixture of compositions: а) without sludge; b) with 20 % of sludge;
c) with 40 % of sludge; 1 – without ammonium dihydrogen phosphate; 2–4 – with 1, 2, and 3 %
of ammonium dihydrogen phosphate respectively.
A further increase in the mixing duration from 14 to 20 minutes leads to a decrease in the strength of the samples with 1 % of the NH4H2PO4 additive but practically does not affect the strength of the samples with 2 % of this additive. This is explained by a significant slowdown in the hydration of the gypsum binder due to the formation of blocking films of sparingly soluble calcium hydrogen phosphate dihydrate on the surface of its grains, which is especially pronounced at high dosages of the chemical additive.
When a 3 % ammonium dihydrogen phosphate additive is added to the gypsum binder, a significant decrease in the strength of the pressed gypsum composites occurs. This is explained, on the one hand, by the formation of too thick phase films on the surface of binder particles, characterized by increased porosity and reduced elasticity, and, on the other hand, by the fact that when ammonium dihydrogen phosphate is consumed in excess of the optimal amount, an excessive amount of ammonium sulfate (NH4)2SO4 is formed in the molding mixture [11, 21]. This compound is a strong electrolyte that increases the solubility of gypsum and thereby accelerates its crystallization.
When replacing a part of the gypsum binder with a fine filler, mixing must be carried out for a sufficient time to obtain a homogeneous mixture. As our research has shown, with the combined use of sludge additives and ammonium dihydrogen phosphate, in order to produce the material with the highest strength, the mixing time must be carefully chosen.
а)
b)
c)
Figure 2. The dependence of the compressive strength of water-saturated samples
on the duration of mixing of the molding mixture of compositions: a) without sludge;
b) with 20 % of sludge; c) with 40 % of sludge; 1 – without ammonium dihydrogen phosphate;
2–4 – with 1, 2, and 3 % of ammonium dihydrogen phosphate, respectively.
Figs. 1b and 2b show that the composition containing 20 % additives of chemical water treatment sludge and 2 % additives of ammonium dihydrogen phosphate has the maximum strength (Curve 3) with an ideal mixing period of 8 to 12 minutes. With an additive content of 1 % in the mixture (Fig. 1b), the highest strength is achieved after 8–10 minutes of mixing, and at 3 % – after 12–14 minutes. Moreover, an increase in the mixing time of a mixture containing 3 % NH4H2PO4, has a negative effect on the strength of the samples to a much lesser extent, compared with a mixture with 1 % chemical additive. This confirms our assumptions about different hydration rates during the formation of the structure of pressed gypsum composites in the presence of different amounts of film-forming agent additives.
When replacing 40 % of the gypsum binder with sludge (Figs. 1c and 2c, Curve 1), samples without ammonium dihydrogen phosphate additives showed the greatest strength when the mixing duration was 10 minutes. The same mixing time can be called optimal for the composition of the molding mixture with 1 % ammonium dihydrogen phosphate. An increase in the consumption of ammonium dihydrogen phosphate while prolonging the mixing time of the mixture improves the physical and mechanical characteristics of the samples made from them. It is most advisable to mix mixtures with 2 % of the additive for 12–14 minutes, and with 3 % – for 14–16 minutes.
Therefore, the conducted studies have revealed that the duration of mixing gypsum binder with modifying additives is a crucial technological component of the proposed method of modifying the structure of pressed gypsum composites to improve their strength and water resistance. The highest strength (45–60 MPa) and softening coefficient (0.65–0.70) were demonstrated by samples of composite binders containing 20–40 % of sludge and 2 % of ammonium dihydrogen phosphate with an optimal mixing duration of 10–14 minutes.
It has been established that the determination of optimal dosages of additives and the duration of mixing of the molding mixture must be carried out considering the combined influence of these factors on the physical and mechanical properties of the material. Therefore, at the second stage of the research, we conducted experimental and statistical studies, the purpose of which was to obtain mathematical models of the physical and mechanical characteristics of the proposed material, allowing us to obtain products with specified properties. For this purpose, a complete factorial experiment was planned and implemented using the B3 Box composition plan [27], and the studied prescription and technological factors and rational intervals of their variation were adopted based on the above analysis of the performed pre-tests.
The conditions for planning and conducting the experiment are presented in Table 2.
Table 2. Conditions of the experiment.
Code
Code value
The factors under study and the intervals of their variation
X1 – chemical water treatment sludge
additive, %
X2 – ammonium
dihydrogen phosphate additive, %
X3 – duration of mixing, min
Main level хi 0
0
20
2
10
Variation range Δхi
∆
10
1
5
Upper level хi max
+1
30
3
15
Lower level хi min
−1
10
1
5
The effect of the water-solids ratio (W/S) and pressing pressure (P) on the studied physical and mechanical properties of gypsum composites was not considered in the paper. Based on the results of the previous studies, these factors were stabilized for all points of the experiment (V/S = 0.19; P = 40 MPa).
During the experiment, six control samples were produced for each point of the plan. The changes in the compressive strength of dry and water-saturated samples ( and respectively), the average density of the material its water absorption by mass , and softening coefficient were studied.
The average values of the experimental data for each point of the experiment, as well as the calculated theoretical values of the studied technical characteristics of gypsum modified composites are presented in Table 3.
Table 3. Test results of control samples and calculated values of physical and mechanical properties of gypsum modified composites.
Experiment number
Experiment plan
Property under study
X1
X2
X3
Y1 – strength in the dry state, MPa
Y2 – strength in the water-saturated state, MPa
Y3 – average density, kg/m3
Y4 – water absorption by mass, %
Y5 – softening coefficient
1
+1
+1
+1
47.3 / 48.1
30.5 / 30.6
1920 / 1918
7.0 / 7.1
0.63 / 0.64
2
+1
+1
–1
41.0 / 39.9
24.8 / 25.5
1921 / 1923
7.1 / 7.2
0.62 / 0.62
3
–1
+1
+1
40.5 / 39.4
22.4 / 23.6
1928 / 1926
8.2 / 8.2
0.58 / 0.59
4
–1
+1
–1
28.3 / 30.1
17.9 / 17.4
1933 / 1932
8.0 / 8.0
0.56 / 0.57
5
0
+1
0
54.7 / 54.4
38.3 / 37.1
1948 / 1950
6.0 / 5.9
0.69 / 0.68
6
0
0
+1
58.2 / 59.4
42.1 / 41.3
1941 / 1945
6.0 / 6.2
0.71 / 0.70
7
+1
0
0
54.3 / 55.5
39.1 / 37.9
1926 / 1923
6.1 / 6.0
0.68 / 0.69
8
0
0
0
63.6 / 62.6
45.1 / 44.4
1948 /1949
5.8 / 5.4
0.71 / 0.72
9
–1
0
0
48.9 / 48.2
31.5 / 32.0
1930 / 1932
6.6 / 6.8
0.66 / 0.65
10
0
0
–1
58.2 / 57.5
39.5 / 40.6
1950 /1946
6.2 /6.2
0.70 / 0.70
11
0
–1
0
54.2 / 55.0
34.1 / 35.7
1950 / 1947
6.2 / 6.6
0.64 / 0.65
12
+1
–1
+1
41.8 / 39.9
22.3 / 22.7
1920 / 1921
7.9 / 7.9
0.58 / 0.57
13
+1
–1
–1
44.3 / 45.4
28.1 / 27.1
1915 / 1917
8.2 / 8.2
0.60 / 0.59
14
–1
–1
+1
34.0 / 35.0
19.6 / 18.8
1930 / 1928
8.8 / 8.7
0.55 / 0.55
15
–1
–1
–1
40.5 / 39.6
22.6 / 22.4
1925 / 1926
8.7 / 8.6
0.57 / 0.57
16
0
0
0
63.4 / 62.6
46.3 / 44.4
1948 /1949
5.7 / 5.4
0.73 / 0.72
17
0
0
0
65.2 / 62.6
46.9 / 44.4
1948 /1949
5.8 / 5.4
0.72 / 0.72
Note: before the slash – an experimental result, after the slash – a theoretical result calculated based on the experimental statistical models obtained.
The obtained experimental and statistical models of the most important technical characteristics of gypsum modified pressed composites allow for a comprehensive analysis of their relationship with the factors studied in the accepted ranges of their variation. This makes it possible to set the conditions for obtaining pressed gypsum products with specified properties. The results of the experimental statistical modeling are in good agreement with the results of the above-described pre-tests and therefore are not reanalyzed further in this paper. Considering the goals of the study, the fifth three-factor quadratic model is of the greatest relevance since it describes how the modifying filler, the amount of NH4H2PO4 additive, and the duration of mixing the gypsum mixture impact the softening coefficient (water resistance) of pressed composites:
(1)
As a result of the transformation of the equation (1), three quasi-one-factor models can be obtained that make it possible to evaluate the individual effect of the three studied indicators on the water resistance of gypsum modified pressed composites:
(2)
(3)
(4)
Quasi-one-factor models of the effect of the carbonate filler consumption the dosage of the chemical additive , and the duration of mixing of the molding mixture are graphically presented in Fig. 3.
Figure 3. Quasi-one-factor models of the effect of additives of chemical water treatment
sludge (Х1) and ammonium dihydrogen phosphate (Х2), as well as the duration of mixing (Х3)
on the softening coefficient (water resistance) of the material: “+” – upper level;
“–” – lower level.
As shown in Fig. 3, the modifying effect of the carbonate filler has a significant impact on increasing the softening coefficient of pressed gypsum composites. Moreover, the highest values of the softening coefficient can be achieved if the other two factors studied ( and ) are at the upper level. If the dosage of the ammonium dihydrogen phosphate additive and the duration of mixing are at the lower level, then the softening coefficient, on the contrary, decreases with an increase in the consumption of the sludge additive from chemical water treatment of thermal power plants.
The effect of the addition of ammonium dihydrogen phosphate on the water resistance of pressed gypsum composites is of a different nature. The material has the highest softening coefficient with the lowest consumption of ammonium dihydrogen phosphate additive, if the other two factors are at the lower level. If = +1 and = +1, then the softening coefficient increases with the increasing dosage of the NH4H2PO4 additive.
The duration of mixing the molding mixture, in comparison with the dosages of the carbonate filler and the chemical additive, has a less significant effect on the water resistance of the material. Fig. 3 clearly shows that the optimum is always located in the experiment area. In this regard, when looking for a combination of factors which will result in the material with the maximum softening coefficient, that is, when solving the optimization problem of the first type, using the method described in [25], the optimal control function for the duration of mixing the mixture can be calculated. This function will look like:
(5)
It follows from the equation (5) that each of the consumptions of carbonate filler and ammonium dihydrogen phosphate has its own optimal mixing time.
After substituting (5) into (1), the adjustment equation of the softening coefficient of the material is obtained for
(6)
The adjustment equation shows how the softening coefficient changes depending on factors and if the duration of mixing the mixture is optimal for any combination of dosages of chemical water treatment sludge and ammonium dihydrogen phosphate additives. This equation can be used to calculate the appropriate additive dosages for a given softening coefficient and to determine the rational mixing time for the molding mixture.
4.Conslusion
As a result of the research conducted, it is possible to draw the following conclusions:
It is shown that it is possible to increase the strength and water resistance of gypsum pressed composites by replacing a significant part of the gypsum binder with a finely dispersed carbonate filler and creating protective shielding films from sparingly soluble calcium hydrogen phosphate dihydrate on the surface of forming microcrystals of gypsum and filler due to the modifying effect of the ammonium dihydrogen phosphate additive.
It has been found that in the proposed technology, along with the dosages of modifying additives, the physical and mechanical properties of the material are significantly affected by the duration of mixing of the molding mixture, during which chemical interaction occurs between its components, which increases the strength and water resistance of gypsum modified pressed composites.
Making use of mathematical experiment planning techniques, experimental and statistical models of the most important technical characteristics of the proposed material have been developed, which make it possible to determine the conditions for obtaining pressed gypsum products with specified properties.
An adjustment equation is proposed that makes it possible to establish rational dosages of modifying additives and the optimal duration of mixing of the molding mixture for a given softening coefficient.
10.34910/MCE.142.2
691.311
gypsum composites; strength; water resistance; structure formation; experimental and statistical modeling
https://engstroy.spbstu.ru/article/2026.142.2/
RAR
RUS
14203-14203
Department of Civil Engineering, University of Diyala
Abd
Noor
noorazoz29@gmail.com
Diyala, Iraq
Department of Civil Engineering, University of Diyala
Abbas
Jasim
Jasimalshamary@yahoo.com
Diyala, Iraq
Experimental model of shallow foundation over treated expansive soil using fly ash-based geopolymer
In fact, constructing low-weight buildings over expansive soil is usually risky, directly influencing urban development. Therefore, this type of soil needs to improve before loading, and one of the new and sustainable methods is using geopolymer materials. This study uses fly ash-based geopolymer with different percentages (i.e., 0.5, 1, 2, 4, and 6 %). After that, the 2 % was selected and applied directly to the laboratory model to improve the surface layers of the soil. It can seem that, the soil transfers from high swelling potential to very low when using this percentage in both the free surface of soil and soil under load. The free swelling decreased from 13 to 0.3 %, at the 6 % geopolymer ratio, and the swelling pressure also reduced from 230 to 7 kPa at the same geopolymer ratio.
1.Introduction
In general, expansive soils are usually widely distributed all over the world [1]. This type of soil is very sensitive to changes in its moisture content and due to these large changes in volume occur [2]. Because of that, possibly inflicts considerable damage to the structures, especially for lightweight buildings and pavements, etc. [3] This performance is commonly due to containing montmorillonite in clay minerals. In practice situations, the swelling problems stabilized using various materials and additives to improve its engineering properties [4, 5]. One of the new and sustainable materials used as a new generation for soil improvement is geopolymer, which combines industrial waste materials including for example fly ash [6, 7].
A number of previous studies included the use of fly ash with different activators for soil stabilization. Bose [8] studied the soil stabilization use of fly ash with different percentages (0, 20, 40, 60, 80, and 90 %), the addition of fly ash can reduce the free swell index, and values of swelling pressure as well as change the grain size of treated soils, also reducing the plasticity of expansive soil. Hasan [9] found that adding 20 % of fly ash can reduce free swelling by 33 % and swelling pressure by 40 %. This fly ash percentage reduced the expansive soil's liquid limit, while the plastic limit decreased by 18 %, and the plasticity index was 28 %. Phani Kumar & Radhey [10] used the same proportions of fly ash that were already used by Hasan [9], but the potential of swelling of the soil was less, where it was found that both swelling pressure and free swelling were reduced by 50 % when 20 % of fly ash was used. As for the liquidity limit, it decreased by 12.5 %, and the plasticity limit increased by 36 %. Salim [11] added different percentages of fly ash (5, 10, and 15 %) of the dry weight to expansive soil containing different percentages of bentonite (30, 50, and 70 %), where the optimum percentage was 5 % of fly ash, which results in less swelling pressure and less swelling with better workability. Das & Parhi [12] used alkaline activated fly ash (AAFA) in different proportions (0, 5, 10, and 15 %) of the dry weight of the soil, to improve the properties of the expansive soil, where it was observed that the free swelling index (FSI) decreased to one-third of the initial soil FSI at 15 % AAFA content. Sharma & Sivapullaiah [13] used a mixture of fly ash and ground granulated blast furnace slag (GGBS) in different proportions to improve the properties of the expansive soil, including the potential swelling, FSI, and swelling pressure. The results revealed that the swelling behavior decreases with increasing the concentration of the binder and adding 1 % lime to the binder, increasing the ability to reduce swelling. Phanikumar [14] compared the addition of two pozzolanic materials, lime, and fly ash, at different percentages (0, 2, 4, and 6 % and 0, 10, and 20 %, respectively), as the results indicated that their effects on the swelling potential and swelling pressure are close. Radhakrishnan et al. [15] studied the effect of mixing chlorides (Aluminum Chloride AlCl3, Magnesium Chloride MgCl2) with fly ash in different proportions (0, 0.5, 1, 1.5, and 2 % and 0, 5, 10, 15, and 20 %, respectively), on the swelling properties (swell potential, FSI, and swell pressure) on expansive soil, the effect of 1 % AlCl3 with 10 % fly ash was found to be more effective than the other two. Sabat & Pradhan [16] discussed the suitability of mixtures of expansive soil reinforced with fibers and stabilized with fly ash as subgrade materials for flexible pavement. The results showed that 1 % of propylene fibers with a length of 12 mm is the optimal ratio for strengthening expanded soil stabilized with an ideal amount of 20 % fly ash, as the swelling pressure decreased to 17 kN/m2.
Fly ashes are microscopic particles made up primarily of silica, alumina, and iron [17]; it is a pozzolanic substance produced by coal burning in thermal power stations [18]. It is simply a non-plastic fine silt whose composition changes depending on the type of coal burned. Currently, the production of fly ash much outnumbers its consumption.
Based on previous studies, it can seem that the application of treated soil with geopolymers in laboratory models is very limited with different thicknesses. Therefore, this study involves optimizing layers of different thicknesses of expansive soil using a small-scale laboratory model.
2.Methods and Materials
2.1.Expansive Soil
The expansive soil was prepared by mixing 80 % bentonite and 20 % sand, with adding the optimum water content, and mixing it with the soil until the mixture becomes homogeneous. Table 1 displays the expansive soil's mechanical and physical characteristics.
Table 1. Summary of the engineering properties of the soil used.
Standard
Soil Property
Value
ASTM D 4318
USCS classification
CH
ASTM D 854
Specific gravity
2.78
ASTM D-422
Liquid limit, %
98
Plastic limit, %
45
Plasticity index, %
53
ASTM D-3084
Free swelling, %
13
Swelling pressure, kPa
230
Compression index (Cc)
0.16
ASTM D-1557
Maximum dry unit weight, kN/m³
14.45
Optimum moisture content, %
26
ASTM D-2216
Unconfined compressive strength(qu), kPa
245
2.2.Fly Ash and Alkali Activator
In this study, fly ash class (F). Its chemical properties are shown in Table 2, sodium hydroxide, and sodium silicate were used as alkaline activators to accelerate the reaction. Sodium hydroxide (NaOH) is a chemical compound that dissolves in water and is used in wide fields. Sodium silicate (Na2SiO3) is an important chemical compound also known as water glass.
Table 2. Fly Ash chemical composition.
Chemical Composition
Percent (%)
SiO2
47.67
Al2O3
27.73
Fe2O3
18.42
CaO
5.11
MgO
2.65
TiO2
1.3
K2O
0.6
2.3.Experimental Model
In this study, a container used with dimensions 450 × 450 × 500 mm made of iron with a thickness of 40 mm, shown in Fig. 1. It is supplied with water by an elevated small tank. The clay soil thickness is 20 cm, and the upper soil layers were treated starting from the soil surface divided into three stages, the first stage is 2 cm the second is 3 cm, and the last stage is 4 cm. A square aluminum plate 10 × 10 cm with a thickness of 1 cm was used as a foundation and placed in the center of the container. This foundation is designed to carry 30 kPa. After applying the surface load, the soil is supplied with water until the soil reaches a fully saturated state, and then a dial gauge reading is taken with respect to load and free soil surface every 24 hours.
Figure 1. Experimental model.
3.Result and Discussion
In this part, the results include the basic tests (i.e., Atterberg limits, swelling pressure, and free swelling) that were carried out for the expansive soil before and after improvement. When adding the fly ash-based geopolymer at five different percentages (0.5, 1, 2, 4, and 6 %) from the weight of dry soil. Based on the results, it can be concluded that the best percentage was 2 %, which shows high improvement and means transferring the soil stat from high to very low according to potential expansion (PE) (ASTM-D4829). Depending on that, this percentage has been selected as an additive in the small-scale laboratory model.
Fundamental Tests Results
Fig. 2 shows the effect of fly ash-based geopolymer percentages on the Atterberg limits, where it was observed that they gradually decrease with the increase of the fly ash-based Geopolymer content. The liquid limit value decreased from 98 to 65 %, i.e., a decrease of 34 %, the plasticity index value also decreased from 53 to 35 % at 4 % geopolymer. The interpretation of the above results is that, as a result of replacing coarse fly ash particles with fine soil particles, the addition of fly ash-based geopolymer reduces the volume fraction of clay in the soil, induces the flocculation of the clay particles, and raises the number of coarse particles [19], decreasing the soil's liquid limit and plasticity index.
Figure 2. The effect of fly ash-based geopolymer percentages on the Atterberg limits.
By performing swell-consolidation tests with samples that were 75 mm in diameter, and 16 mm in thickness, the swelling potential and swelling pressure of the blended samples were calculated at various percentages of fly ash-based geopolymer, according to ASTM-D4546, the free swell method. With a rise in geopolymer content, swelling potential and swelling pressure dropped and reached more than 90 % as illustrated in Fig. 3. As a result of reduced clay minerals' capacity to absorb water due to ion exchange with geopolymer, pozzolan, which fills gaps, is produced.
Figure 3. Effect of geopolymer on swelling percentage and swelling pressure.
Results of Laboratory Model
In this part, the results obtained from the laboratory model have been presented and discussed. Fig. 4 shows the amount of swelling under the foundation and the amount of swelling in the free surface for the untreated soil over a 16-day time period. It was observed that the swelling of the free surface of the soil reached 26 mm, which means 13 % of the total height of expansive soil. While the amount of swelling under load was 10.25 mm, which represents 5 %. According to the ASTM-D4829, the swelling of the free surface of the soil can be categorized as high PE. On the other hand, soil swelling under load is about two and a half times less than the free swelling of the soil surface. The differences between the swelling potential (i.e., free and under load) are a result of the influence of load that usually reduces the swelling potential.
Figure 4. Swelling of free soil and load vs. time for untreated soil.
Fig. 5 shows the free swelling of the untreated soil surface and the free swelling of the treated soil surface in the form of layers of different thicknesses starting from the soil surface to a depth of 4 cm. The free swelling of the treated soil surface decreases at 2, 3, and 4 cm depth by 33.5, 61, and 75 %, respectively, of the free swelling of the untreated soil surface, according to ASTM-D4829, the swelling of the free surface of the soil treated at 2 cm can be categorized as medium PE, and at 3, 4 cm – as low PE.
Figure 5. Swelling vs. time for treated and untreated soil for free surface.
Fig. 6 shows swelling under load for the untreated soil and swelling under load for the treated soil. The amount of swelling under load was less when treating the soil at a depth of 2, 3, and 4 cm by 40, 70, and 88.5 %, respectively, than swelling under load for the untreated soil, according to ASTM-D4829, the swelling of the load of the soil treated at 2 cm can be categorized as low PE and at 3, 4 cm – as very low PE.
Figure 6. Swelling vs. time for treated and untreated soil for load.
Fig. 7 shows the relationship between free swelling of the soil surface and swelling under load with the improved soil layers. It can seem that the thickness of improved soil increased the swelling decreased in all cases. The reason for this is that swelling occurs more in the surface layers compared to the lower layers because the weight of the soil itself prevents swelling with depth. In addition, these improvements occur due to the activity of the geopolymer as it forms a pozzolanic material that fills the gaps and causes flocculation of the clay particles by the exchange cation [20].
Figure 7. Prediction of swelling with an Increase in thickness of the improved layer.
4.Conclusion
Based on the results, it can seem that the use of geopolymer material is good results as expansive soil improvement, the following conclusion is developed:
The best percentage for improving soil properties such as liquid limit, plastic limit, and plasticity index was 4 %, both the liquid limit and plasticity index decreased by 34 %, and the liquid limit decreased by 33 %, but it is possible to use 2 % because the difference in percentage improvement was convergent.
The addition of 4 % reduced free swelling and swell pressure by 96 and 95 %, respectively. While the 2 % can reduce free swelling and swell pressure to 90 and 92 %, respectively. That is, it is possible to use 2 % geopolymer instead of 4 % because the results of free swelling and swelling pressure at those percentages of geopolymer are very close.
For free soil, the soil transfers from high swelling potential to a low state at 3 and 4 cm improved layer. While, for the soil under load, the soil transfers from a high to low state at 2 cm, and to very low at both 3 and 4 cm.
10.34910/MCE.142.3
624
expansive soil
fly ash
geopolymer
swelling potential
swell pressure
https://engstroy.spbstu.ru/article/2026.142.3/
RAR
RUS
14204-14204
57194440967
0000-0003-4153-1046
Don State Technical University
Nesvetaev
Grigory
nesgrin@yandex.ru
Rostov-on-Don, Russian Federation
56056531000
0000-0002-9133-8546
Don State Technical University
Chepurnenko
Anton
anton_chepurnenk@mail.ru
Rostov-on-Don, Russia
57196034514
0000-0002-2341-9811
Don State Technical University
Koryanova
Yulia
koryanova.yi@mail.ru
Rostov-on-Don, Russian Federation
54950122700
0000-0002-5205-1446
Don State Technical University
Yazyev
Batyr
ps62@yandex.ru
Rostov-on-Don, Russia
Autogenous shrinkage and stress-strain state of massive foundation slab
Introduction: The reliability of the assessment of the level of tensile stresses in order to predict the risk of early cracking during the construction of massive monolithic reinforced concrete structures increases when taking into account the deformations of autogenous shrinkage, which are often unreasonably ignored. Purpose of the study: to quantify the effect of autogenous shrinkage of concrete on the formation of a stress-strain state in the early period of the construction of massive monolithic reinforced concrete structures. Materials and methods: Modeling the formation of a stress field without taking into account relaxation in a massive block of 20×20×2 m with a layer overlap time ("layer birth time") of 4 hours. At the first stage of research, stress calculated as a result of the development of only temperature deformations. At the second stage of research, stress calculated as a result of the development of temperature deformations and autogenous shrinkage. When calculating the temperature field, the layered laying of the concrete mixture is taken into account by assigning an abnormally high coefficient of thermal conductivity to each layer before it is laid (1000 W/(m·°C)) and zero heat capacity. Results. The effect of a decrease in the heat transfer coefficient from 23 to 3 W/m2·°C on top of the plate is resulted in a decrease in the maximum stress level in the section by 28–32 % when taking into account only temperature deformations and by 14–27 % when taking into account temperature deformations and autogenous shrinkage. The influence of the kinetics of heat dissipation and hardening during the transition from the rapid group to the slow group is resulted in a decrease in the maximum stress level in the cross section to 6 % when taking into account only temperature deformations. Possible increase in the maximum stress level in the cross section up to 20–56 % take place if temperature deformations and autogenous shrinkage are taking into account.
1.Introduction
The experience of building foundations for unique structures [1, 2] shows that the multiplicity of factors determining the quality of massive monolithic foundations determines the urgency of careful study of the totality of issues related to the technology of their concreting, taking into account prescription factors, technical capabilities, and weather conditions [3–5]. To a large extent, the risk of early cracking in the initial period of concrete hardening in massive monolithic reinforced concrete structures is due to temperature and shrinkage deformations of hardening concrete [6–9]. The reliability of the assessment of the level of tensile stresses increases when shrinkage deformations of concrete are taken into account [10, 11], especially when the overlap time of the layers is more than 4 hours and when using concrete mixtures from different plants (on different cements). In some cases, when calculating the stress level in the early period of construction, shrinkage deformations are ignored [11, 12], which is an unjustified simplification. For example, according to [13], in concretes with a compressive strength of more than 50 MPa without taking into account temperature deformations at the age of 3–5 days, the values of autogenous shrinkage-induced (AS-induced) stress may reach 3 MPa with a deformation of AS up to 0.1 mm/m. Obviously, it is impractical to ignore such a factor.
Stress increments in the assessment of the stress-strain state of massive monolithic foundation slabs can be determined by the equation [11, 14]:
(1)
where is the E-modulus of concrete, is the average increment of total deformation over the thickness of the slab, is a coefficient of thermal expansion, is the difference between the temperature at the point at the current and previous time steps, is the increment of shrinkage deformation, is the increment of creep deformation, is the Poisson's ratio of concrete. The multiplier in the denominator in Equation (1) takes into account the work of concrete under conditions of biaxial tension (compression) and the value is determined by the equation:
(2)
where is the thickness of the foundation plate.
Equations (1, 2) are applicable in the case where the base limits the deformations of the slab along z but does not limit the deformations of the slab along x and y (A soil with a low modulus of elasticity compared to concrete).
Thus, in the early period of concrete hardening, AS poses a certain danger to early cracking of structures [10, 15]. As is well known, AS is caused by a change in the volume of cement paste (stone) during hydration [16, 17]. The share of AS in the total volume change of cement paster during hydration depends on many factors [18]. AS may affect the stress level of massive structures in the early period [19] and increases with a decrease in the W/C ratio, depends on the properties of cement and the presence of chemical admixtures [16, 17], for example, superplasticizers, develops most intensively in the first 5–7 days of hardening [20–22]. AS-induced stress is also influenced by the properties and concentration of aggregates [23].
The multiplicity of factors that influence on the included in Equations (1, 2) values, depending on prescription and technological factors, determines the expediency of using numerical methods in calculating the stress level [24–26]. In this regard, the relevance of obtaining the dependences of the quantities included in Equations (1, 2) on prescription and technological factors in order to model the stress-strain state in the early period of the construction of massive structures is obvious. Self-compacting concrete and concrete mixtures of classes S4, S5 are widely used in the construction of massive monolithic reinforced concrete structures, including foundation slabs. The kinetics of hardening, the formation of the structure and properties of modern concretes with organomineral modifiers has a number of features in comparison with concretes from traditional mixtures [27, 28]. It is well known that to ensure the solidity of massive structures, an important task is to regulate temperature and shrinkage deformations [15, 29–30]. In this regard, in order to predict more precisely the risk of early cracking, it is relevant to assess the role of AS in the formation of a stress-strain state in the early period.
2.Materials and Methods
The research was carried out by modeling the formation of a stress field in a massive block of 20×20×2 m:
at the first stage, studies only of temperature deformations without taking into account stress relaxation;
at the second stage of studies both of temperature and shrinkage deformations without stress relaxation were used.
Stress relaxation was not taken into account, which in some cases is accepted in modeling [31], because, firstly, the task was to assess the effect on the stress level of AS, and, secondly, data on stress relaxation in the early period are few and ambiguous [32].
In the simulation, 8 layers of the laid concrete mix of 0.25 m each were assumed with a layer overlap time ("layer birth time") of 4 hours. The simulation uses the results of laboratory studies and the authors' experience in regulating temperature stresses during the construction of foundation slabs with a volume of 1560–1640 m3 in 14–29 hours using concretes grades B25 and B40 from highly mobile concrete mixtures. When modeling, the following provisions are adopted:
the stress increment is calculated according to Equations (1, 2);
the calculation of temperature fields is based on the solution of the differential equation of thermal conductivity [7, 12]:
(3)
where is the coefficient of thermal conductivity; is the temperature; is the density of internal heat sources; is the density of concrete; is the specific heat of concrete; is the time.
And since the temperature distribution over the cross section, with the exception of the peripheral zones of the foundation slab, is one-dimensional, Equation (4) is used to determine the function instead of Equation (3) [12, 33]:
(4)
The boundary conditions in the case of convective heat exchange with the environment have the form:
(5)
where is the normal to the surface, is the heat transfer coefficient, is the temperature of the environment.
Sequential laying of concrete mix layers after 4 hours when calculating the temperature field is taken into account by assigning an abnormally high coefficient of thermal conductivity to each layer before it is laid (1000 W/(m·°C)) and zero heat capacity, which is mathematically equivalent to the absence of a layer [34]. The temperature field was calculated using the finite element method in the program developed by the authors in the MATLAB environment. The influence of temperature and degree of hydration of concrete on the thermal characteristics of the slab was not taken into account. The entry in Equations (4, 5) means that changes from an abnormally high value in the actual absence of a layer to a standard value after its "birth."
The kinetics of concrete compressive strength was given by the authors equation similar equation in EN 1992-1-1:
(6)
where was assumed to be 0.2, 0.25, and 0.38, respectively, for rapid (R), moderate (M), and slow (S) hardening concretes; = 0.25 days; is the actual curing time, taking into account the "maturity of concrete," days ( = 20 °C).
The function of the strength of concrete under axial tension was given by the authors equation:
(7)
The assumption is made that the modulus of elasticity of concrete (E-modulus) is equal under compression and tension. The E-modulus function is given by authors Equation (8) similar equation in EN 1992-1-1 obtained by processing experimental data of 65 pairs of " " values with a range of compressive strength " " from 9.3 to 69.2 MPa and E-modulus from 21.4 to 37.75 GPa of concretes made from self-compacting mixtures and mixtures classes S4, S5:
(8)
Time of hardening of the concretes ranged from 1 to 28 days. The concretes were made using three Portland cements of class C42.5, of which one is additive-free (CEM I), two with various mineral additives (CEM II). In experimental studies, concrete mixtures were used both without additives and with superplasticizing admixtures PCE or NF type 1 and type 2. The concretes hardened under normal conditions at = 20...22 °C, > 95 %, or in the laboratory at = 22...25 °C, ≈ 45...55 %, or in the massive volume at a peak temperature of 39...42 °C. The values of " " in real time were recalculated to the actual time according to the "concrete maturity" indicator, taking into account the actual temperature conditions [32, 35].
AS was determined within 7 days using samples made from cement paste with a W/C = 0.27 [21]. Samples after demoulding were isolated from water evaporation. Six cements CEM I and CEM II with three different superplasticizing admixture PCE and NF – type 1 and 2 were used.
The properties of the studied cements are presented in Table 1.
Table 1. Studied cements.
Cement properties
Cements
1
2
3
5
6
7
CEM I 42.5Н1,6
CEM II/А-Ш 42.5Н1,6 slag
CEM I 42.5Н2,6
CEM I 42.5Н3,6
CEM II/А-П 42.5Н СС4,7
CEM I 42.5 Н5,6
Flexural/Compressive strength, N/mm2,
2 days age
4.59/22.4*
4.84/23.4
5.04/28.5
5.82/30.1
5.09/23.5
6.49/32.3
Flexural/Compressive strength, N/mm2,
28 days age
7.81/51.5
8.58/55.3
8.55/50.9
8.96/55.6
9.95/54.6
8.79/63.1
Normal consistency of cement paste, %
22.75
25.0
26.5
26.25
27.25
27.75
Setting time, min
115
150
150
170
150
140
C3S
66.5+0.2
66.7+0.3
64.9+0.4
68.4+0.4
63.8+0.3
58.3+0.3
C2S
12.1+0.3
11.9+0.3
12.1+0.3
10.7+0.3
16.1+0.3
15.3+0.4
C3A
6.8+0.2
6.9+0.3
5.3+0.3
4.96+0.2
4.4+0.2
7.7+0.2
C4AF
12.1+0.2
12.3+0.2
15.1+0.2
13.1+0.3
14.2+0.2
12.3+0.2
SO3
2.37+0.2
2.11+0.2
2.84+0.2
2.78+0.2
2.82+0.2
3.06+0.3
Q7, kJ/kg
340.9
298.1
310.6
321.1
358.8
332.4
1–5 – manufacturers; 6 – GOST 31108-2020; 7 – GOST 22266-2013; * – flexural/compressive strength
Information about the admixture used is presented in Table 2.
Table 2. Admixtures.
No.
Admixture
Standard
Chemical basis
1
"POLIPLAST SP-4"
TS 5745-026-58042865-2007
sopolymers based on naphthalene sulfonic acid (NF type 1)
2
"LINAMIX PC" type 2
TS 5745-033-58042865-2008
polyoxyethylene derivatives of polycarboxylic acids and polyethylene glycol (PCE)
3
PFM-NLK
TS 5745-022-58042865-2007
mixture of sodium salts of polymethylene naphthalene sulfonic acids,
hydrophobizing and air-entraining component (NF type 2)
The kinetics of AS of the studied cements is described by the authors equation [21]:
(9)
where = 0.35, 0.45, 0.6; = 0.6, 0.75, 0.9, respectively, for the moderate (M), slow (S), and very slow (VS) groups; = 0,25 days; is the actual curing time, taking into account the "maturity of concrete," days.
At the age of 3 days, the average value of for the M group was 0.75, this is the upper curve; for the S group, it was 0.6, this is about the average; for the VS group, it was 0.4, this is the lower curve (Fig. 1).
Figure 1. Kinetics of autogenous shrinkage of cements: 0, 1, 2, 3 – without admixtures,
with the admixtures of PCE, NF type 1 and type 2, respectively; C1–C7 – cements according
to Table 1; M – moderate, S – slow, and VS – very slow groups.
At the age of 3 days, 14 of the 24 formulations had a value of more than 0.6 and 10 less than 0.6. The AS values of the studied cements with admixture at the age of 7 days were changed from 0.361 to 0.864 mm/m. The predicted AS values of concretes with cement consumption from 260 (C20/25) to 420 kg/m3 (C35/45) at the age of 7 days ranged from 0.004 to 0.15 mm/m.
The AS function, depending on the compressive strength was given by the authors equation:
(10)
where: – relative AS; – relative compressive strength of concrete; – the coefficients is equal minus 0.64 and 2.48 for the VS group, 0.24 and 1.4 for the S group, 1.4 and minus 0.12 for the M group, respectively.
According to [17], deformations of AS equal 0.15 mm/m in concretes with a value of W/C = 0.5 occurred after 15 hours of hardening. According to [32], the AS of concretes with a strength of 40...50 MPa at the age of 7 days was equal 0.025...0.07 mm/m. According to [36], already at the age of one day, the AS value of concrete with a strength of 32 MPa was 0.05 mm/m, and concrete with a strength of 100 MPa was 0.1 mm/m. According to [37], at the age of 5 days, concretes with a strength of 54.8...58.1 MPa had an AS value of 0.125...0.175 mm/m. According to [38], at the age of 5 days, concretes with a strength of 41...64 MPa had an AS value of 0...0.195 mm/m. According to [39–41], at the age of 3 days, the AS value of concrete changes from 0.025 to 1 mm/m. An analysis of a not complete database on AS shows that AS values can vary over a very wide range. In our calculations, the AS value of concretes in 7 days age was carefully assumed to be 0.025 mm/m for concretes of class C20/25 and 0.25 mm/m for concretes of class C35/45.
The heat dissipation function was given by the authors equation:
(11)
where are the coefficients; – is the duration of the induction period (is a parameter adjusted to take into account different setting times due to different mix temperatures caused by the use of retarding/accelerating admixtures).
The values of the coefficients in Equation (11) are assumed to be = 0.14, 0.19, 0.24; = 0.4, 0.51, 0.62, respectively, for concrete R, M, S groups. In Equation (6), the values of the coefficient are assumed to be 0.2, 0.35, 0.5, respectively. The value of is assumed according to experimental data from 110 to 190 MJ/m3. The heat transfer coefficient from the upper surface is assumed from 3 to
23 W/(m2 °С). The lower surface of the block is assumed to be in contact with a 3 m thick soil mass with a density of 1800 kg/m3 and a specific heat capacity of 750 J/(kg °C) and a thermal conductivity coefficient of 0.9 W/(m °С).
3.Results and Discussion
According to [9], in a 2 m thick concrete structure of class B35 with a module of surface less than 1.1, the maximum temperature in the center of the structure reached 64.5 °C after 75 hours, and 29 °C on the surface after 52 hours. According to [39], in a 2 m thick concrete structure with a cement consumption of 340 kg/m3, the maximum temperature in the center of the structure reached 63.8 °C after 72 hours, and 28.2 °C on the surface after 48 hours. In our research, when modeling the stress-strain state of a 2 m thick foundation slab made of concrete classes from B25 to B45 with a specific heat dissipation of concrete from 110 to 190 MJ/m3, depending on the kinetics of heat dissipation and the conditions of heat exchange "top surface – ambient," the maximum temperature in the center was from 43.3 to 77.2 °C after a time of 59 to 204 hours. On the top surface, respectively, from 26.1 to 60.2 °C after 24–161 hours. When measured in a real foundation slab with a thickness of 2 m and a volume of 1642 m3 made of rapid-hardening concrete of class B25 with a specific heat dissipation of 130...140 MJ/m3, the maximum temperature in the center after 45 hours was 67.2 °C, on the top surface covered with a tarpaulin, after 25 hours a temperature of 55 °C was recorded at an ambient temperature from 18 to 24 °C.
In structures, the value of maximum tensile stress depends on the concrete properties and structural parameters. For example, according to [40], the excess of the stresses of the tensile strength of concrete was recorded after about 84 hours. According to [6], for concrete with a design strength of 80 MPa, the tensile stress exceeded the tensile strength of 3 MPa after about 54 hours. According to [39], on the top surface of the slab, after 72 hours, at the moment of maximum temperature in the center of the slab, the value of the tensile stresses reached 2.43 MPa. In our studies, in the real foundation slab described above, the tensile stresses on the upper surface reached a value comparable to the ultimate tensile strength of 2.32 MPa after 110 hours with a center-top surface temperature difference of 26 °C.
A significant amount of research is devoted to the role of shrinkage in the formation of the stress-strain state of structures. For example, in the study [41] of the stress-strain state of massive structures as a function of temperature and shrinkage deformations, ACI 209 and CEB-FIP formulas were compared using thick-walled cylinders as example. Tensile stress values from shrinkage to 7 MPa are obtained. According to [4], when modeling with AS according to various standards at a maximum temperature in the center of the structure of 47.7 °C and a maximum temperature difference "center-top" of 28.7 °C, the values of tensile stresses from 1.8 to 4.17 MPa were obtained. According to our data, when modeling the stress-strain state described above foundation slab with a thickness of 2 m, the maximum values of tensile stresses on the upper surface of the foundation slab ranged from 1.59 to 5.4 MPa. However, the authors are not aware of studies that would provide an assessment of the effect of AS, taking into account its magnitude and kinetics on the stress-strain state, which are presented in this paper.
Fig. 2 shows the dependence of the calculated values of the maximum tensile stress level in the section of the massive block under consideration on the class and kinetics of concrete hardening, the magnitude and kinetics of concrete heat dissipation, and the heat transfer coefficient taking into account AS deformations of concrete.
Figure 2. Dependence of the maximum stress level in the section on the kinetics and magnitude
of concrete heat dissipation, concrete class, and heat transfer coefficient from the upper surface: 25, 35, 45 – concrete class C20/25, B35 (В35 in accordance with Russian standard (similar C30/37)), C35/45 respectively; R, S – groups according to Equation (6);
3, 23 – heat transfer coefficient.
Fig. 3 shows the dependence of the calculated values of the maximum tensile stress level in the section of the massive block under consideration on the class and kinetics of concrete hardening, the magnitude and kinetics of concrete heat dissipation, and the heat transfer coefficient without taking into account AS deformations of concrete.
Figure 3. Dependence of the maximum stress level in the section on the kinetics and magnitude
of concrete heat dissipation, concrete class, and heat transfer coefficient from the upper surface: 25, 35, 45 – concrete class C20/25, B35 (В35 in accordance with Russian standard (similar C30/37)), C35/45 respectively; R, S – groups according to Equation (6);
3, 23 – heat transfer coefficient.
Fig. 4 shows the ratio of the calculated values of the maximum stress level in the section over a 7-day period, taking into account AS deformation and without AS deformation of concrete depending on the class and kinetics of concrete hardening, magnitude and kinetics heat dissipation of concrete, heat transfer coefficient.
Figure 4. The ratio of the maximum stress levels in the section from the kinetics and magnitude
of the heat dissipation of concrete, the concrete class, and the coefficient of heat transfer
from the upper surface: 25, 35, 45 – concrete class C20/25, B35 (В35 in accordance
with Russian standard (similar C30/37)), C35/45 respectively; R, S – groups
according to Equation (6); 3, 23 – heat transfer coefficient.
Analysis of the simulation results shows that:
the effect of a decrease in the heat transfer coefficient from 23 to 3 W/m2·°C is manifested in a decrease in the maximum stress level in the section, depending on the values of the studied factors, by 28–32 % when taking into account only temperature deformations and by 14–27 % when taking into account temperature and AS deformations of concrete;
the influence of the kinetics of heat dissipation and kinetic of strength during the transition from the R group to the S group is manifested in a decrease in the maximum stress level in the section, depending on the values of the studied factors, to 6 %, when taking into account only temperature deformations, and in an increase in the maximum stress level in the section by 20–56 % when taking into account for temperature and AS deformations;
as a result, when taking into account temperature and AS deformations of concrete together at the "layer birth time" (layer overlap time) of 4 hours, taking into account all the values of the studied factors, manifests itself in a predicted decrease in the stress level to 30 % or the level increase up to 36 %.
Obviously, it is impractical to ignore such a possible influence of AS deformation on the maximum stress level in the section. The ambiguity of the influence of AS deformation of concrete on the formation of the stress field is due to the manifestation of a possible inconsistency between the kinetics of heat dissipation of cement (concrete), as a result of which a temperature field and temperature deformations are formed, and the kinetics of AS deformation of concrete (Fig. 5).
Figure 5. Diagram of the possible development of the resulting deformation depending both
on the kinetic of temperature and kinetic of AS deformation of concrete: R, S – rapid,
slow groups for the kinetics of heat dissipation and strength of concrete respectively;
M, S, VS – medium, slow, very slow groups for AS deformation of concrete, respectively;
ε(T), ε(AS) – temperature and AS deformations of concrete, respectively.
Obviously, with a rapid increase in temperature deformations and a slow development of AS deformation, compensation for temperature deformations practically does not occur at the stage of temperature rise. At the stage of cooling, AS deformation increases the total deformations. Such a variant of the influence of AS deformation can be considered as negative. With a slower increase in temperature deformation and a faster development of AS deformation, it partially compensates for temperature deformations at the stage of temperature rise. This variant of the influence of AS, depending on the ratio of the magnitudes of temperature and shrinkage deformations, as well as the values of the E-modulus and the ultimate strength of concrete under axial tension, may be positive. From the results shown in Fig. 6, it follows that for the studied combinations of cements with admixtures, there is no single dependence of the kinetics of AS deformation on the kinetics of heat dissipation
Figure 6. Dependence of relative AS deformation on relative heat dissipation: 1, 3 – at the age
of 1 and 3 days, respectively, C0 – without superplasticizing admixtures.
Table 3 shows the maximum and minimum values of and at diurnal age.
Table 3. Maximum and minimum values of Qτ/Q7 and εAS,τ/ εAS,7 at diurnal age.
Indicator
Cement / admixture
max = 0.742
0.181
–
0.244
3 / NF type 2
min = 0.452
0.265
–
0.586
6 / NF type 2
max = 0.43
–
0.659
0.653
6 / NF type 1
min = 0.061
–
0.722
0.084
5 / PCE
From the results presented in Table 3, it follows that, under the condition of a faster development of AS deformation of concrete relative to the kinetics of heat dissipation, preference should be given to C6 cement in combination with an NF type 2 or NF type 1 admixture. The least preferred option is C5 cement in combination with a PCE admixture.
4.Conclusions
It is shown that ignoring the residual deformation in concrete when calculating the stress-strain state of massive monolithic structures during early hardening can lead to an error in the results, depending on the structural parameters and properties of concrete, by more than 30 %, both in terms of reducing and increasing the tensile stress in concrete. For the studied combinations of cements with admixtures, there is no single dependence of the kinetics of AS deformation of concrete on the kinetics of heat dissipation. For concretes intended for the construction of massive monolithic reinforced concrete structures the value for example, at the age of one day, can be used as one of the criteria for choosing a rational combination of the "cement + admixture." Of course, the values and should also be considered as very important criteria, depending on the construction parameters and concrete properties.
10.34910/MCE.142.4
693.547
massive monolithic structures
overlap time of layers
risk of early cracking
heat dissipation of concrete
temperature deformations
autogenous shrinkage
strength kinetics
stress level
https://engstroy.spbstu.ru/article/2026.142.4/
RAR
RUS
14205-14205
Civil Engineering Department, University of Technology
Ghaddar
Maha
maha.g.ghaddar@uotechnology.edu.iq
Baghdad, Iraq
57205235403
0000-0001-8332-8607
Civil Engineering Department, University of Baghdad
Al Shamaa
Mushriq Fuad Kadhim
mushriqf@coeng.uobaghdad.edu.iq
Baghdad, Iraq
Civil Engineering Department, University of Baghdad
Hussain
Ahmad
dr.ahmadalshimmeri@coeng.uobaghdad.edu.iq
Baghdad, Iraq
0000-0003-1304-0303
University of Baghdad
Karkush
Mahdi
mahdi_karkush@coeng.uobaghdad.edu.iq
Baghdad, Iraq
Department of Reconstructions & Projects, University of Baghdad
Jabur
Maher
maher.murad@coeng.uobaghdad.edu.iq
Baghdad, Iraq
Numerical approach for improving and enhancing the behavior of I-shaped damper using diagonal stiffeners added to the web of the damper
Comprehensive researches to consider the behaviour of the metallic dampers have proven that the I-shaped link, designed to act as shear mechanism, serves as a ductile fuse. However, a key limitation is the reduced strength and stiffness of the system when these links are directly connected to the diagonal member of a concentrically braced frame. To improve the structural performance of I-shaped links, researchers have suggested increasing the web plate thickness, which, in turn, raises the load on the brace members. A novel solution to address this issue involves incorporating diagonal stiffeners into the dampers. This study numerically and parametrically investigates the effect of stiffening I-shaped links using finite element analysis. Numerical studies show that the stiffeners not only prevent the web plate from buckling (which improves its performance) but also share the imposed loading. By including stiffeners, a thinner web plate can be utilized instead of a thicker one to achieve greater ultimate strength and stiffness. The parametric analysis highlights that the thickness of the stiffeners plays a more critical role than the properties of the flange or web plate in determining damper performance. Based on these results, an optimal configuration for I-shaped dampers is proposed in this study.
1.Introduction
The concentrically braced frames (CBFs) have accepted system with high lateral strength and stiffness but suffered from low ductility. To address these shortcomings, researchers have explored various strategies, with passive energy dissipation devices, specifically metallic dampers, emerging as an effective solution to enhance CBF performance. Researchers have discovered various approaches to overcome these shortcomings, with metallic dampers, emerging as an effective solution to enhance CBF performance.
Various dampers have been presented to enhance the CBF’s behavior. The diagonal components of CBFs tend to buckle under compressive forces, which dwindled their ability to dissipate energy. This buckling issue is a significant limitation regarding the CBFs performance. Utilizing shear damper proposed by Ghamari et al. [1], this problem was solved. Their design featured a shear plate enclosed within a cylindrical casing, offering improved ductility and energy dissipation. The shear damper was more economical and easier to produce in comparing with other famuse damper such as ADAS [2], TADAS [3], shear panels [4], slit dampers [5], and buckling-restrained braces (BRBs) [6]. Expanding on this concept, Jeong, Ghamari et al. [7, 8] suggested an octagonal cover instead of the cylindrical one, which showed effective performance under shear, flexural, and combined shear-flexural conditions. However, the octagonal configuration increased the damper’s weight, making manufacturing and installation more challenging. Drawing inspiration from the effective shear link behavior in eccentrically braced frames (EBFs) [9–19], researchers proposed integrating an I-shaped shear element directly into CBFs. Experimental studies [20–31] confirmed that I-shaped dampers enhance CBF performance by improving ductility, preventing brace buckling, and simplifying fabrication and installation compared to other dampers. These studies showed that I-shaped dampers increase overstrength and ductility but often reduce the system’s overall stiffness and strength. The influence of the I-shaped damper’s flange on system behavior was also evaluated.
Despite their benefits, I-shaped dampers, like other directly attached CBF dampers, tend to compromise the system’s ultimate strength and stiffness. Two common approaches to address this issue include: a) using an I-shaped section with a thicker web or b) increasing the number of dampers and braces. However, thickening the damper web has a limited impact on stiffness and can elevate stresses in surrounding brace elements, potentially leading to buckling. In such cases, the damper may fail to perform effectively, resulting in reduced strength, stiffness, and brittle behavior. Similarly, adding more dampers to restore stiffness introduces higher construction costs and architectural constraints.
A promising alternative is the incorporation of stiffeners into the I-shaped damper design. Stiffeners require minimal material and are straightforward to implement, as illustrated in Fig. 1. This idea successfully compensates for lacks in stiffness and ultimate strength, offering a practical solution to enhance the seismic behavior of CBFs.
Figure 1. Strengthening of I-shaped damper with stiffeners.
2.Methods and Materials
To ensure that the damper yields prior to other structural components, the elements outside the damper must be designed to withstand forces exceeding the damper’s capacity. This force denoted as which is calculated using the following expression:
(1)
where is the ratio of ultimate stress of the materials to the yielding stress. Also, represents the overstrength factor, a critical parameter explored in this study, and denotes the nominal strength of the damper. The overstrength factor, is determined by evaluating the ratio of the nominal strength to the strength, at which the first plastic hinge forms, as expressed in:
(2)
According to the AISC 341-16 [32] standard for shear links with I-shaped cross-sections, the shear strength is calculated as:
(3)
where is the yield stress of the web plate, and and represent the width and thickness of the web plate, respectively. The standard specifies that a shear-dominated mechanism governs when the ratio defined as:
(4)
where is less than 1.6. For dampers equipped with X-shaped stiffeners, the nominal strength is computed as:
(5)
where is the number of web plates, is the shear strength contributed by the X-shaped stiffeners, and is the shear strength of the flange plates. The flange plate’s contribution to shear strength is derived from the formation of flexural hinges at both ends of the flange plates. Using plastic theory, the maximum strength of the flange plate is achieved when the plastic moment is reached, and is calculated as:
(6)
where is the effective height of the damper, and the plastic moment is determined using plate-shell theory as:
(7)
Here, and represent the width, thickness, and yield stress of the flange plate, respectively.
At the onset of hinge formation, the strength accounts for contributions from the web plate, flange plate, and stiffeners, and is expressed as:
(8)
where is web plate stiffness that calculated as Accordingly, is the plastic deformation, and and are the stiffness contributions of the flange plates and stiffeners, respectively. This formulation ensures a comprehensive assessment of the damper’s behavior under seismic loading, enabling robust and resilient structural performance.
3.Results and Discussion
Numerical Models
In this study, a numerical analysis is conducted to evaluate the performance of a metallic damper, with a focus on the influence of the flange plate and its properties on the damper’s behavior. The numerical studies are performed using the finite element (FE) method with simulations performed using ANSYS. The damper’s components were modeled using the SHELL181 element, which is well-suited for capturing large displacements and buckling while ensuring convergence during nonlinear simulations.
The FE models were defined with consistent geometric parameters: a web plate thickness of 6 mm, damper height of 140 mm, stiffener width of 160 mm, and flange width of 160 mm. The variables examined include the web plate thickness and stiffener thickness which result in variations in shear strength and stiffness ratio A36 steel, characterized by a yield strength of 235 MPa and a Young’s modulus of 200 GPa, was used for all damper components.
Lateral loading was applied to the damper using simplified boundary conditions, as depicted in Fig. 2. Given the pinned connections between the beam and columns, the main frame is assumed to have no influence on the damper’s behavior, and vice versa. Since the primary objective of this study is to assess the damper’s performance and validate proposed analytical relationships, only the damper itself is analyzed. The interaction between the damper and a diagonally braced frame will be explored in future research.
Figure 2. Boundary conditions of simplified damper.
Validation of Finite Element Results
According to Fig. 3, the FE results were validated against the experimental results of the I-shaped shear damper detailed in [33] through FE simulation. As can be seen from Fig. 3, the FE simulation is relatively compatible with the experimental results on predicting the hysteresis curve and the yielding state of the damper.
a)
b)
c)
Figure 3. Comparing the experimental test and FE results in a) hysteresis curve;
b) the tested damper [33]; c) simulated the damper.
4.Discussion
Comparing the Curves of Stiffened/Unstiffened Dampers
As discussed in Section 1, the shear capacity, stiffness, and overstrength of the damper are closely tied to the properties of the web plates. As a result, the adoption of low-yield-point (LYP) steel is expected to significantly enhance the damper’s performance under shear loading. To explore this further, Fig. 4 illustrates both the load-rotation relationship and the stiffness versus rotation behavior, plotted for a range of ρ values to provide a comprehensive view of the damper’s response under varying conditions.
The load-rotation graph serves as a valuable tool for evaluating the overall performance of different damper configurations. As shown in the figure, the addition of stiffeners to create a reinforced I-shaped damper shifts the curve upward, indicating improvements in key mechanical properties. Specifically, this modification leads to a noticeable increase in elastic stiffness, ultimate strength, and energy dissipation capacity, all of which are critical for effective seismic performance. These enhancements suggest that the stiffened damper can better withstand cyclic loading, making it a promising option for structures in earthquake-prone regions.
To further contextualize these findings, the results suggest that design decisions regarding stiffener placement and material selection should prioritize the intended loading conditions. For applications where elastic stiffness is paramount, such as in structures requiring minimal deformation under moderate loads, the reinforced I-shaped damper offers clear advantages. However, in scenarios where the damper is expected to undergo significant nonlinear deformation, alternative approaches such as optimizing the web plate geometry or exploring hybrid material combinations may be necessary to achieve the desired performance. These insights provide a foundation for tailoring damper designs to specific structural demands, balancing cost, complexity, and seismic resilience.
Figure 4. Comparing the curves of the I-shaped and stiffened dampers.
Stiffness
Fig. 5 illustrate the stiffness of the I-shaped and stiffened dampers. Referring to Figure 5, either damper with and without stiffeners have no degradation in stiffness. Also, adding stiffeners improve the stiffness considerably. By increasing the thickness of the stiffeners, the stiffness has raised. So, the stiffeners affect the stiffness in elastic and inelastic zones.
Figure 5. Comparing the stiffness of the I-shaped and stiffened dampers.
Overstrength
Given that elements external to the damper must be designed to withstand amplified forces based on the overstrength factor the values of for various dampers are presented Fig. 6. Also, the of stiffened dampers are divided by the I-shaped dampers in Fig. 6 to consider the effect of the X-stiffeners on of the damper. Results indicated that have > 1.5, which exceeds the AISC-recommended value of 1.5. Consequently, it is advised to adopt an of 2.0 for damper design since its average is 1.97.
Figure 6. The Ω of the dampers with X-stiffeners.
Referring to Fig. 7, the values for dampers with different flange plate thicknesses are normalized against those with a flange thickness of 10 mm. The data suggest that thicker flange plates result in higher values. Similarly, the seventh column compares the of stiffened dampers to unstiffened I-shaped dampers with identical With the exception of the damper at = 0.25, the addition of stiffeners increases by 2 % to 19 %.
To better evaluate the influence of various parameters on overstrength, Fig. 5 plots against different variables. The results indicate that increasing the ratios and leads to a higher though this relationship is nonlinear. Conversely, increasing the slenderness of stiffeners, as represented by the ratios and tends to decrease To mitigate significant reductions in it is recommended to limit to 35 and to 50.
Figure 7. Comparing the Ω of the I-shaped and stiffened dampers.
Limitation of Stiffeners Slenderness to Achieve Suitable Overstrength
To evaluate the appropriateness of the proposed constraints of = 40 and = 60, Fig. 8 presents a detailed analysis by plotting ultimate strength and stiffness against these variables. Consistent with the behavior observed for overstrength, the data reveal that both ultimate strength and stiffness diminish as the slenderness of the stiffeners, represented by and increases. This trend highlights the sensitivity of the damper’s performance to geometric configurations, particularly in terms of how slender stiffeners can compromise structural efficiency.
At the specific thresholds of = 40 and = 60, the graphs indicate a notable shift: the rate of reduction in both stiffness and ultimate strength begins to stabilize, transitioning to a more gradual decline. This leveling-off suggests that these limits mark a practical boundary where further increases in slenderness yield diminishing negative impacts on performance. Consequently, adopting = 40 and = 60 as design constraints appears to be a well-founded and robust recommendation, balancing structural integrity with practical implementation.
These findings have broader implications for damper design optimization. By adhering to the proposed limits, engineers can ensure that the damper maintains adequate overstrength that represent the stiffness, strength, and economic aspect under anticipated loading conditions, particularly in seismic applications where energy dissipation is critical. Moreover, the data in Fig. 6 underscore the importance of carefully calibrating stiffener geometry to avoid excessive slenderness, which could lead to premature degradation of performance. To further validate these limits, future studies could explore their applicability across a wider range of damper configurations and loading scenarios, potentially refining the thresholds for specific use cases. For now, these constraints offer a reliable guideline for achieving efficient and resilient damper designs.
Figure 8. The ultimate strength and stiffness versus slenderness.
5.Conclusions
This study numerically and parametrically explored the performance of an I-shaped damper enhanced with stiffeners, confirming its effective behavior. The key findings are outlined below:
The stiffened I-shaped damper functions as a ductile fuse, efficiently absorbing energy imposed on structures during loading.
Incorporating stiffeners into the I-shaped damper significantly improves its elastic stiffness, ultimate strength, and energy dissipation capacity. However, their impact on stiffness in the nonlinear deformation phase remains minimal.
The average overstrength factor, for both stiffened and unstiffened I-shaped dampers was calculated as 1.97, surpassing the AISC-recommended value of 1.5. As a result, a design value of = 2.0 is suggested for practical applications.
The characteristics of the flange plates and stiffeners influence Thicker flange plates lead to an increase in Additionally, with the exception of the damper at = 0.25, stiffeners contribute to a 2 % to 19 % rise in
Increasing the slenderness ratios of stiffeners ( and ) results in a reduction of To counteract significant declines, it is recommended to limit to 35 and to 50.
At the proposed limits of = 35 and = 50, the decline in stiffness and strength stabilizes, indicating a more gradual reduction. These thresholds are therefore considered a reliable and effective design guideline.
10.34910/MCE.142.5
624
CBF
passive damper
seismic
overstrength
enhancing
https://engstroy.spbstu.ru/article/2026.142.5/
RAR
RUS
14206-14206
0009-0005-3149-7341
Peter the Great St. Petersburg Polytechnic University
Sorokin
Vladislav
vladislavs2018@mail.ru
St. Peterburg, Russian Federation
Protection of buildings and structures in the zone of influence of metro tunnel construction
This article proposes a constructive solution for a protective combined geotechnical barrier to minimize the impact on existing buildings and structures from the construction of a subway tunnel. The barrier is utilized during the construction of a metro tunnel in weak soils on a developed area. The object of research is existing buildings and structures in the area affected by underground construction. The research subject is the additional deformations of the foundations and structures of existing buildings and structures. The aim of the study is to reduce the impact on neighboring buildings during the construction of the metro tunnel. Method. The research was conducted using numerical modeling. Three-dimensional and planar finite element models were developed in the Plaxis software package for analysis. Results. The values of additional foundation settlements and stresses in the building structures were determined. The effectiveness of the installation of a geotechnical barrier along the escalator tunnel is confirmed. The barrier reduces the maximum absolute foundation settlements of the surrounding buildings by an average of 40 % and the relative differential settlement by 66 %.
1.Introduction
The object of research is existing buildings and structures in the area affected by underground construction. The research subject is the additional deformations of the foundations and structures of existing buildings and structures.
Underground construction in dense urban areas causes additional deformations of foundations and structures of neighboring buildings. This issue is particularly relevant in the construction of metro systems in urban environments due to the risks of damaging surrounding buildings and the significant economic costs involved in their restoration. For example, the construction of the world's first London Underground was accompanied by emergency situations. In May 1861, during earthworks on Euston Road, sidewalks, gardens, telegraph lines, water pipes, and building facades were damaged [1]. The impact of tunneling on existing structures is discussed in works [2–6]. Protection of surrounding buildings from the effects of underground construction is addressed in works [7–10].
Large difficulties arise during the construction of tunnels in weak clay water-saturated soils. According to research [11], urban development on weak clay soils is heavily influenced by foundation construction technologies due to the possibility of disrupting the soil structure. This occurs as a result of changes in the stress-strain state in the soil mass during the excavation of a deep pit or tunneling. Disturbances from construction equipment are also transmitted to the soil. Vibrational and impact actions lead to compaction or vibratory loosening of soils. This results in the vertical deformation (technological settlements) of neighboring building foundations.
Weak clay soils have long-term creep. Creep is not studied during standard engineering and geological surveys. Accordingly, this property is not taken into account when calculating foundation settlements. At the same time, buildings constructed on such soils receive settlements during their existence that exceed permissible values. Additional foundation settlements and forces in the structures of existing buildings depend on the work technology, equipment, type of enclosing structures, etc.
Therefore, to obtain accurate values for additional deformations, it is necessary to consider the various impacts that occurred on the soil in the surrounding area prior to construction.
The dependency of the excavation locations relative to buildings on their stress-strain state was examined in works [12].
Damage to existing buildings during underground construction primarily results from the engineering-geological and hydrogeological conditions of the construction site, the construction technology employed, and the complexity of the project [13]. Consequently, ensuring the safety of buildings and structures within the zone of influence of underground construction is a paramount task for engineers. Research in the field of structural safety assessment is described in works [16–20].
The relevance of this work lies in the development of an optimal design for a protective structure (barrier). This barrier is intended to minimize the force factors that arise in the structures of existing buildings as a result of metro tunnel construction. Particular attention is paid to material efficiency. The study utilizes nonlinear models for the behavior of soil and building materials. It is expected that the research results will serve as a guideline for engineers involved in geotechnical calculations when assessing the impact of excavation in constrained urban environments.
The aim of the study is to develop a design for a protective screen to reduce the impact on adjacent buildings during the construction of a metro tunnel.
To achieve the stated objective, it is essential to accomplish the following tasks:
to analyze the influence of the excavation's location on the stress-strain state of the buildings' primary load-bearing structures;
to determine the additional foundation deformations and the resulting forces in the load-bearing structures during tunnel excavation by performing a coupled analysis of the "soil-foundation-structure" system;
to develop a methodology to mitigate the impact of tunnel excavation on existing buildings.
2.Materials and Methods
During construction (reconstruction) in densely built-up conditions, the safety of existing buildings and structures is affected by the following man-made impacts:
1) temporary change in the groundwater regime – change in hydrogeological conditions in the construction area due to local water decrease;
2) temporary technological impacts associated with the method of work.
The safety criteria for existing buildings are restrictions on settlements, their relative differences or tilts by limiting values:
(1)
(2)
where – the magnitude of additional deformations (their differences) of surrounding structures due to the influence of the -th group of factors on the foundation; – is the minimum distance between the foundations of a neighboring building; – are deformations from previous impacts in an unstabilized state of the foundation. This value is determined by calculation based on the results of assessing the degree of stabilization of settlements when monitoring the position of sensors on cracks.
The value of the maximum permissible additional settlement, relative difference or tilt of adjacent structures should be calculated for a specific building or structure in their actual deformed state, when performing a joint calculation with the foundation [19]. The obtained values guarantee the safety of existing buildings' structures during the development of additional deformations.
The study examines the construction of an escalator metro tunnel as part of an underground complex. The underground space is located in a historic urban area. For such buildings, strict requirements are imposed to limit deformations of their foundations [22–30]. A general view of the underground structure is shown in Fig. 1.
Figure 1. General view of the underground complex (the tunnel is not shown conditionally).
The physical and mechanical characteristics of the soils are presented in Table 1. The foundations of existing buildings are water-saturated sands. The groundwater level is located at a depth of 1.5–2.0 m from the day surface.
Table 1. Physical and mechanical properties of soils.
No.
Ground
γ, kN/m3
ν
С, kPa
φ, °
E50ref = Eoedref (Eurref), MPa
1
Made ground
18
0.30
–
–
–
2
Saturated sand
20
0.32
1
25
13.55
3
Silty clay of fluid consistency
19.3
0.37
13
16
2.4
4
Clay loam of fluid consistency
18.2
0.37
38
9
3.1
5
Plastic sandy loam
21.3
0.35
18
24
7.5
6
Clay loam of low plasticity
20.8
0.35
33
29
9.0
7
Semi-solid loam
21.5
0.35
31
19
12.4
8
Dislocated hard clays
21.1
0.2
69
22
12.7
9
Hard clays
21.8
0.2
273
5.3
30
The engineering and geological conditions of the construction site are shown in Fig. 2.
Figure 2. Engineering and geological conditions of the construction site.
The surrounding buildings consist of load-bearing longitudinal and transverse walls, which are united in height by horizontal floor discs. The walls, 0.55–2.0 m thick, are made of brick with lime-sand mortar. The ceilings are supported by metal or wooden beams. Some are vaulted from brick and concrete. The foundations are made of strip rubble stone on a lime-sand mortar. Roofs of complex shape are built on wooden rafter structures.
The calculation of deformations of the foundations of surrounding buildings was carried out using the finite element method in the Plaxis 3D software package [29]. The computational area has plan dimensions of 350×360 m. The total depth of the soil foundation is 70 m. The load-bearing soil layer model (Ground No. 9 in Table 1) is the Hardening Soil model. For soft soils, the Mohr–Coulomb model was used. The deformation and strength characteristics of the soils of building foundations are taken into account the results of dynamic sounding. These characteristics reflect the actual state of the soil. This approach involves changing properties. During the entire operational cycle, they changed under the influence of loads from the building and various man-made factors. The surrounding buildings are specified by volumetric elements with a Linear Elastic material model to take into account their rigidity in the formation of the stress-strain state of soils. The specific gravity of buildings is assumed to be 5 kN/m3 [30], which is an average for most buildings. The work of the soil in contact with structures is modeled using reducing interface coefficients These coefficients reflect a decrease in the strength characteristics of soils due to technological influences. The values of these coefficients depend on the soil, the material of the enclosing structure, the method of its construction, etc. and are determined from reference data. For concrete and reinforced concrete structures = 0.67 [31]. The design diagram of the underground complex and the metro tunnel is shown in Fig. 3.
Figure 3. General view of a three-dimensional design diagram in Paxis 3D.
The calculation of additional settlements was carried out in the following steps:
1) calculation of natural stresses of the soil mass;
2) zeroing out movements, activating surrounding buildings;
3) activation of the fencing of the pit and interfaces, changing the hydrological conditions of the soil within the pit;
4) stage-by-stage development of the pit;
5) stage-by-stage excavation of the metro tunnel.
The following options for enclosing structures of the starting pit were considered:
1) diaphragm wall made of reinforced concrete 0.6 m thick, made using trench technology;
2) a row of drilled tangential piles with a diameter of 0.6 m with injection of pile mates for waterproofness of the fence;
3) soil-cement barrier using Jet-Grouting technology.
Options for pit enclosing structures are shown in Fig. 4.
Figure 4. Options for pit enclosing structures.
The stiffnesses of the considered options are presented in Table 2. When determining the moments of inertia per linear meter of the fence for Option No. 2, areas of jet grouting of joints are not taken into account, and for Option No. 3 the conditional width of the fence is assumed to be 0.6 m.
Table 2. Rigidity characteristics of fencing design options.
Option No.
Material
Modulus of elasticity E, MPa
Sectional area A, m2
Axial stiffness ЕA, kN
Bending stiffness ЕJ, кNm2
1
Reinforced concrete
30000
0.60
1.8×107
5.4×105
2
Reinforced concrete
30000
0.45
1.35×107
3.0×105
3
Soil-cement
150
≈0.6
9×104
2.7×103
The calculation results for the pit fencing options are presented in Table 3.
Table 3. Calculation results for pit fencing options.
Option No
Maximum displacement
Umax, mm
Maximum bending moment Мmax, kNm/m
Maximum normal force in strut rail Nmax, kN/m
1
23
1050
1334
2
25
735
1387
3
56
192
2434
Figure 5. Comparison of pit enclosing wall options.
An analysis of enclosing structure options is presented in Fig. 5. For Option No. 1, soil displacements and forces in strut rail structures are minimal. This is due to the high stiffness of the structure. Also, Options No. 2 and No. 3 do not guarantee the tightness of the pit during its development. The entry of water into the pit entails the development of suffusion processes and additional settlements of buildings. Therefore, the enclosing structure of the pit is made diaphragm wall of reinforced concrete 0.6 m thick, made using trench technology.
The excavation of the escalator tunnel at an angle of 30 ° is carried out by an earth pressure balance tunnel boring machine with a starting pit with a diameter of 23 m. The fencing of the pit is made of reinforced concrete walls and transverse diaphragms for the railway tracks of the overhead crane. The lining of the productivity tunnel is made of prefabricated reinforced concrete blocks measuring 1.0 m and 0.4 m thick. The outer diameter of the tunnel is 10.5 m. The depth of the tunnel is about 65 m from the ground surface level. A general view of the starting pit and escalator tunnel is shown in Fig. 6.
Figure 6. General view of the starting pit and escalator tunnel.
3.Results and Discussion
Additional Settlements of Building Foundations
When Tunneling a Metro Tunnel Without Protective Measures
The results of calculations of vertical and horizontal soil displacements without protective measures are shown in Fig. 7.
Figure 7. Isofields of soil mass deformations in the zone of influence
of tunnel excavation without protective measures.
The foundation diagram of existing buildings that fall within the influence zone of tunnel excavation is shown in Fig. 8. The area of 130×60 m is taken as the boundary of the influence zone.
Figure 8. The foundation diagram of foundations of existing buildings and structures
in the zone of influence of tunnel excavation.
To determine additional deformations of building structural elements along design sections 1-1 and 2-2, two-dimensional models were compiled in Plaxis 2D. The structures were modeled using rod elements. In the places where walls and ceilings meet, hinges are introduced in the design diagram. Additional structural deformations of existing buildings are shown in Figs. 9 and 10.
Figure 9. Diagrams of additional settlements of neighboring buildings (section 1-1 in Fig. 7). Maximum value Sad,max = –5.943 mm, (ΔS/L)max = 0.0006.
Figure 10. Diagrams of additional settlements of neighboring buildings (section 2-2 in Fig. 7). Maximum value Sad,max = –4.634 mm, (ΔS/L)max = 0.0005.
The limiting value of the additional settlement of the foundations of neighboring buildings is =
= 5 mm, and the relative difference in settlement = 0.0004. Thus, the deformations of neighboring buildings exceed the permissible limits. Protective measures must be taken to ensure safety.
Additional Settlements of Building Foundations
When Digging a Metro Tunnel with a Geotechnical Barrier
The protective barrier consists of diaphragm wall of reinforced concrete 0.6 m thick and 27–35 m deep. Different depths are made in order to save materials. The diaphragm walls are located parallel to the tunnel axis at a distance of 3 m from it. They limit horizontal displacements of the soil mass. The barrier also includes two diaphragm slabs 4 m thick. The diaphragms are made using single-component Jet-Grouting technology. This technology is discussed in [32–34]. Soil-cement slabs increase the rigidity of the structure in the transverse direction to the axis of the tunnel and prevent it from floating. A general view of the geotechnical barrier is shown in Fig. 11.
Figure 11. General view of the combined geotechnical barrier.
The results of calculations of vertical and horizontal soil displacements using a barrier are shown in Fig. 12.
Figure 12. Isofields of soil mass deformations in the zone influenced by tunneling during the construction of a geotechnical barrier.
Additional structural deformations of existing buildings are shown in Figs. 13 and 14.
Figure 13. Diagrams of additional settlements of neighboring buildings (section 1-1 in Fig. 7). Maximum value Sad,max = –3.469 mm, (ΔS/L)max = 0.0002.
Figure 14. Diagrams of additional settlements of neighboring buildings (section 2-2 in Fig. 7). Maximum value Sad,max = –2.831 mm, (ΔS/L)max = 0.0002.
According to the calculation results, additional settlements and their relative differences do not exceed the maximum permissible values. The use of a geotechnical barrier ensures the safety of existing buildings.
4.Conclusions
A significant amount of destruction of existing buildings occurs during construction and reconstruction near them [35, 36]. In dense urban areas, additional deformations of building foundations are subject to strict requirements. It is necessary to ensure safe work and eliminate the risk of emergency situations.
The total amount of additional settlement is affected by soil compaction processes, changes in hydrogeological conditions in the construction area, and various technogenic impacts during construction and reconstruction. The greatest danger to the structures of the surrounding buildings is represented by the technologies used for the work. Therefore, when designing, it is necessary to adopt the safest technologies. If necessary, protective measures must be developed.
This article examines the effectiveness of using a combined geotechnical barrier when excavating a metro tunnel on soft clay soils. The optimal design solution for the fencing of the starting pit has been determined. Calculations were carried out by numerical modeling in flat and spatial formulations. The effectiveness of the installation of a geotechnical barrier along the escalator tunnel is confirmed. The barrier reduces the maximum absolute foundation settlements of the surrounding buildings by an average of 40 % and the relative differential settlement by 66 %.
Utilizing the Linear Elastic model for adjacent buildings is a limitation that neglects the physical nonlinearity of brick masonry and crack development under differential settlements. However, this approach is justified by the primary focus on the global spatial stiffness of structures for predicting the ground settlement trough, while also reducing the computational cost of the coupled soil-structure interaction analysis in Plaxis 3D.
10.34910/MCE.142.6
69.059.35
surrounding building-up
weak soils
tunneling excavation
zone of influence
geotechnical barrier
base deformation
https://engstroy.spbstu.ru/article/2026.142.6/
RAR
RUS
14207-14207
0000-0001-6244-7965
University of Kufa
Alasadi
Layth Abdulrasool
laitha.alasadi@uokufa.edu.iq
Al Najaf, Iraq
Department of Chemical Engineering, University of Kufa
Al-Yousefi
Hayder
Haydera.Alyousefi@uokufa.edu.iq
Najaf, Iraq
Department of Chemical Engineering, University of Kufa
Kadhem
Ali
alik.hussein@uokufa.edu.iq
Najaf, Iraq
Hydraulic analysis of a combined system: rectangular weir with a rectangular gate
This study explores the hydraulic analysis of a combined system of rectangular sharp-crested weirs and rectangular gates, essential to control water flow in various applications. The weirs and gates are used for flow measurement in water distribution systems, reservoirs, sewage systems, and irrigation networks. The study aims to optimize design and operational strategies for the combined system, contributing to the field of hydraulic engineering. Five models of compound rectangular weirs with rectangular gates were tested using a downstream support. The first three models had a gate width of 10 cm, while the remaining two had a lower opening width of 10 cm and gate widths of 8 cm and 6 cm. The discharge coefficient increases with water head ratio as shown in experimental runs with gate width = 10 cm and lower weir width = 10. The discharge coefficient rises by 10.81 % due to a decrease in the lower opening weir with a constant gate width. It increases by 29.14 % as the gate width decreases with a continuous lower opening.
10.34910/MCE.142.7
69
combined system
weir
gate
hydraulic analysis
https://engstroy.spbstu.ru/article/2026.142.7/
RAR
RUS
14208-14208
0000-0003-4367-0330
Peter the Great St. Petersburg Polytechnic University
Gusev
Konstantin
konstantin0399@yandex.ru
St. Petersburg, Russian Federation
56091980300
0000-0003-3850-424X
Peter the Great St. Petersburg Polytechnic University
Lalin
Vladimir
vllalin@yandex.ru
29 Politechnicheskaya St., St. Petersburg, 195251, Russia
57194431559
0000-0003-4992-2084
Peter the Great Saint Petersburg Polytechnic University
Galyamichev
Alexander
galyamichev@yandex.ru
Polytechnicheskay, 29
0000-0003-4901-4967
Istanbul Okan University
Dogru
Selcuk
seltrue@hotmail.com
Istanbul, Turkey
Bearing capacity of bonded connections in facade structures with metal cassette cladding
Object of the study is a structural adhesive sealant. The aim of the study is to investigate the influence of the hinged facade system (HFS) substructure with metal cassette cladding on the stress–strain state of the adhesive joint by means of numerical analysis (finite element method) and analytical calculations in accordance with international regulatory documents. Method. The bonded connection was investigated using numerical analysis. Four finite element models of the adhesive joint within a HFS were developed, differing in the degree of substructure modeling detail. Static analyses were performed for each modeling approach to determine the resulting shear and tensile stresses in the adhesive joint. The obtained stress values were compared with those calculated according to the provisions of the ETAG 002 guideline. Results. The numerical analysis under isolated modeling conditions demonstrated that when the fastenings of the metal cassette to the HFS substructure are not taken into account, stresses induced by the self-weight of the cladding are uniformly distributed among the adhesive joints, whereas wind-induced stresses exhibit a non-uniform distribution. The distribution of shear stresses within the adhesive joint is significantly affected by the stiffness characteristics of the fastening nodes connecting the cassette to the HFS substructure, with deviations ranging from 28 % to 31 %. Wind-induced stress distribution is influenced by both the fastening system and the adopted structural scheme, resulting in differences of up to 8 % when the load-width participation factor is considered. Neglecting this factor leads to discrepancies of up to 29 % compared with calculations performed in accordance with ETAG 002.
10.34910/MCE.142.8
69.04
stress–strain state
metal cassettes
silicone adhesive sealant
bonded connections
facade structures
numerical modeling
hinged facade systems
https://engstroy.spbstu.ru/article/2026.142.8/
RAR
RUS
14209-14209
Jordan University of Science and Technology
Khedaywi
Taisir
khedaywi@just.edu.jo
Irbid, Jordan
55628523847
0000-0002-8828-2908
United Arab Emirates University
Haddad
Madhar
madhar.haddad@gmail.com
Abu Dhabi, United Arab Emirates
Jordan University of Science and Technology
Kanaan
Osama
osamakanaan100@gmail.com
Irbid, Jordan
Examining impact of ground scrap tire rubber on asphalt mixture properties
Rubber tires present a serious disposal problem as millions of tires are discarded every year producing serious smoke pollution when burnt. The objectives of this research are to examine how recycled rubber from used tires affects properties of mixes of bituminous concrete, and to find optimum rubber content that will provide rubber-asphalt concrete mixtures with best properties. In current study, recycled rubber from used tires is used to modify characteristics of asphalt mixture (AM). Marshall mix design procedure was used to obtain optimum binder content (OBC) and to study impact of rubber on AM properties. Mixtures' moisture susceptibility was determined using Marshall immersion test and indirect tensile strength test. Scrap rubber at 0, 1, 2, and 3 % of total weight of mixture was mixed with crushed limestone or basalt aggregate. By total weight of mixture, asphalt binder (AB) was added at five different percentages (4.0, 4.5, 5.0, 5.5, and 6.0 %). Findings indicate that adding 1 % of rubber to a mixture of limestone and basalt yielded bituminous mixtures with best qualities overall. Increased air voids and flow and reduced unit weight, void in mineral aggregate, and stability has occurred by adding scrap rubber to mixes. Compared to limestone bituminous mixtures, basalt bituminous mixtures exhibit greater stability, unit weight, air voids, voids in mineral aggregate, smaller flow, and slightly smaller voids filled with bitumen. Dry specimens exhibit higher stability compared to wet specimens. Nevertheless, for both limestone and basalt bituminous mixtures at OBC, the flow of wet specimens is higher than that of dry specimens. For both limestone and basalt aggregate-bituminous mixtures, indirect tensile strength reduces as rubber concentration increases at OBC. Also, retained stability ratio (RST) and retained indirect tensile strength ratio (RTSR) in basalt mixtures at OBC is larger than those for limestone aggregate-bituminous mixtures at OBC.
10.34910/MCE.142.9
625.85
bituminous mixtures
ground tire rubber
limestone
basalt
Marshall immersion test
indirect tensile test
moisture susceptibility
https://engstroy.spbstu.ru/article/2026.142.9/
RAR
RUS
14210-14210
55101174500
0000-0001-8650-2375
Peter the Great St. Petersburg Polytechnic University
Sharapov
Dmitry
sharapov.dm@gmail.com
St. Petersburg, Russian Federation
B-6662-2019
57204916380
0000-0002-6877-8420
Peter the Great Saint Petersburg Polytechnic University
Kozinetc
Galina
galina4410@yandex.ru
St. Petersburg, Russia
Peter the Great St. Petersburg Polytechnic University
Kozinetc
Pavel
pavelkozinetc@yandex.ru
St. Petersburg, Russia
Monitoring and reliability assessment of a caisson gate frame in operational conditions
The article explores the comprehensive stress-strain state, reliability assessment, and monitoring of the caisson frame of the S-1 Ship Passage Structure, part of the St. Petersburg Flood Protection Complex. The relevance of this study lies in the uniqueness of the structure, which is unique in the country in terms of its length, as well as the complex and dynamic operating conditions in the Neva Bay. The structure is subject to a variety of factors, including variable hydrostatic pressure, ice loads, and temperature deformations, which shape the stress and strain patterns in the load-bearing elements. Particular attention is paid to the presence of initial residual deformations in the frame, recorded during in-kind inspection, which impact the performance of the structure. The study methodology is based on the finite element method. A detailed finite element model was developed that takes into account the actual geometry of the structure, the physical and mechanical properties of the materials, and the actual initial deformations. A series of calculations was conducted for a set of key operational scenarios: the caisson gate in wet and dry dock, initial moment of movement, and various temperature conditions. For each case, stresses, displacements, and reactions in the support hinge were analyzed. The calculation results showed that local and overall stresses in the frame do not exceed permissible values. Similarly, forces in the support hinge remain below the ultimate load-bearing capacity. However, deformation analysis revealed large deflections, which may be due to installation errors. Large deflections are observed in areas 45-95 meters from the support hinge. Based on the analysis of stress and deformation fields, critical areas of the structure most susceptible to loads and deformations were identified. Recommendations for the placement of vibration, deformation, and strain gauge sensors within the monitoring system were developed for these areas.
10.34910/MCE.142.10
532.517
Caisson gate
stress-strain state
FEM
structural monitoring
reliability
protective structure complex
residual deformations
control and measuring apparatus
https://engstroy.spbstu.ru/article/2026.142.10/