This study investigates the long-term performance of laboratory dam concrete in different curing environments over ten years and the microstructure of 17-year-old laboratory concrete and actual concrete cores drilled from the Three Gorges Dam. The mechanical properties of the laboratory dam concrete, whether cured in natural or standard environments, continued to improve over time. Furthermore, the laboratory dam concrete exhibited good resistance to diffusion and a refined microstructure after 17 years. However, curing and long-term exposure to the local natural environment reduced the frost resistance. Microstructural analyses of the laboratory concrete samples demonstrated that moderate-heat cement and fine fly ash (FA) particles were almost fully hydrated to form compact microstructures consisting of large quantities of homogeneous calcium (alumino)silicate hydrate (C-(A)-S-H) gels and a few crystals. No obvious interfacial transition zones were observed in the microstructure owing to the long-term pozzolanic reaction. This dense and homogenous microstructure was the crucial reason for the excellent long-term performance of the dam concrete. A high FA volume also played a significant role in the microstructural densification and performance growth of dam concrete at a later age. The concrete drilled from the dam surface exhibited a loose microstructure with higher microporosity, indicating that concrete directly exposed to the actual service environment suffered degradation caused by water and wind attacks. In this study, both macro-performance and microstructural analyses revealed that the application of moderate-heat cement and FA resulted in a dense and homogenous microstructure, which ensured the excellent long-term performance of concrete from the Three Gorges Dam after 17 years. Long-term exposure to an actual service environment may lead to microstructural degradation of the concrete surface. Therefore, the retained long-term dam concrete samples need to be further researched to better understand its microstructural evolution and development of its properties.
Chen Lyu, Cheng Yu, Chao Lu, Li Pan, Wenwei Li, Jiaping Liu.
Long-Term Performance and Microstructural Characterization of Dam Concrete in the Three Gorges Project.
Engineering, 2024, 33(2): 258-284 DOI:10.1016/j.eng.2023.04.017
The Yangtze River is the third-largest river in the world and has abundant water and energy reserves. The world-renowned Three Gorges Project (TGP), the largest hydroelectric project in the world, is located in the middle reaches of the Yangtze River [1], [2]. As a typical gravity dam, more than 25 million cubic meters of different types of concrete were used in its construction. The durability damage of concrete, which is the most widely used engineering material in the TGP, hinders the normal operation of the project and can even result in structural failure or dam breakage [3]. Therefore, the long-term durability and service life of concrete in the Three Gorges Dam are particularly noteworthy.
In contrast to the concrete used in other projects, mix proportions of dam concrete commonly have the following remarkable features, namely ① less cementitious material consumption, ② higher volume fraction of aggregate (approximately 90%), ③ larger maximum size of coarse aggregate (up to 150 mm), and ④ local resources (especially for aggregate) [4]. To reduce the risk of thermal cracking, a series of construction technologies, such as cooling water pipes and pre-cooling materials, were adopted in the hydraulic engineering [5], [6]. Moreover, the choice of cementitious material is crucial for controlling the internal temperature increase of the dam. Moderate-heat Portland cement (MHC) and low-heat Portland cement (LHC), as compared to ordinary Portland cement (OPC), essentially alter the phase composition of the clinker to reduce the hydration heat; hence, they are widely used in part or whole mass concrete structures, particularly hydraulic structures [7], [8], [9]. Supplementary cementitious materials (SCMs), such as fly ash (FA) and slag, effectively reduce heat release and improve the performance of concrete [10], [11], [12], [13]. Therefore, SCMs are typically blended with dam concrete as a cement replacement [4]. FA, as compared with other SCMs, reduces the heat release from dam concrete hydration and improves concrete strength and durability. Therefore, it is widely used in hydraulic structures in China and has proven to be an effective solution to the problem of temperature increase in mass concrete [4]. Importantly, FA also contributes to reducing energy consumption and CO2 emissions [14]. However, the use of FA significantly reduces the early performance of concrete. Ternary or quaternary binders (such as FA, limestone powder, silica fume, and other SCMs) and nanomaterials (such as nanosilica and calcium–silicate–hydrate (C-S-H) seeds) have recently been used to increase the early strength of FA concrete (FACB) to minimize the negative effects of FA [15], [16], [17], [18], [19]. The delayed expansion of MgO is an efficient measure to compensate for the long-term thermal shrinkage of dam concrete. Therefore, studies on MgO expansion cement, including MgO expansion agents [20], [21], [22] and high-MgO cement [23], [24], have been widely performed over the past decades [25]. Nevertheless, China is the only country that applies MgO cement to concrete dams on a large scale [4], [25], [26], [27]. During the construction of the TGP, MHC and a large quantity of low-calcium FA were adopted in the dam concrete for the major structure [28], [29]. It should be noted that MHC contains 4.0%–4.5% MgO [4], [30]. Additionally, both a superplasticizer and an air-entraining agent were applied to the TGP dam concrete to facilitate its placement and enhance its durability [31]. The spatial differences in the various positions of the dams lead to differences and complexity in the service environments of the concrete. Therefore, the long-term durability of the dam concrete is of significant concern.
The durability of cementitious composites is often investigated using accelerated degradation tests in the laboratory. Accelerated carbonation has often been used to understand the effects of carbonation on the hydrates [32], pore structure [33] and durability [34], [35], [36] of cementitious composites. Leaching, freezing–thawing, and chemical attacks are frequently accelerated on cement-based materials in the laboratory by curing in high concentrations of chemical solutions or accelerating devices [37], [38], [39], [40]. However, these acceleration methods are performed under certain conditions, and it is difficult to determine the actual long-term performance of dam concrete under service conditions. Therefore, concrete samples drilled from actual projects are frequently investigated to obtain the actual status of the concrete [41], [42], including concrete from dams [43].
Engineers and researchers have successfully used drilled concrete cores to evaluate the service safety of dams. The mechanical properties of cores regularly drilled from Hoover Dam indicated that the strength of low-heat concrete steadily increased over 80 years [44]. The compressive strength and elastic modulus of roller-compacted concrete from the Shapai Dam in China increased by 58% and 14.7%, respectively [45]. Roller-compacted concrete blended with 70% FA from the Yantan Dam exhibited stable hydration products and FA that still had potential reactivity [46]. The sluice pier of the Baisha Reservoir has been in operation for nearly 50 years, and the concrete exhibits evident carbonation and microcracks [47]. The durability of the old concrete from Danjiangkou Dam was evaluated using carbonation and frost resistance data. Microanalysis has also been used to estimate the degree of the alkali-aggregate reaction (AAR) [48], [49]. Zobal et al. [50] reported the carbonation and frost resistance of 55-year-old FACB extracted from the body of Orlík Dam and highlighted the contribution of FA. The old Fengman Dam in China operated with engineering problems for nearly 80 years; however, the aging dam concrete exhibited good mechanical properties and durability [51]. Numerical analyses reproduced the effects of an internal sulfate attack and AAR on dams in Spain [52]. Current and future research on dam concrete should focus on its microstructure [53]. Blanco et al. [54] analyzed the crystalline products and porosity to explain the degradation phenomena of concrete in the 100-year-old Camarasa Dam. Blanco et al. [55] recently summarized the pathologies of four concrete dams using microstructural analysis. Hu [56] concluded that variations in the micropore structures caused by carbonation and calcium leaching induced concrete deterioration in dozens of dams in China. Neumann et al. [57] reported the chemical and mineralogical alterations of the Itaipu Hydroelectric Powerhouse concrete after 40 years of exposure to river water. Rosenqvist et al. [43] comprehensively investigated the chemical and mineralogical characteristics of concrete at different depths and revealed the deterioration mechanism of dam concrete exposed to river water for 55 years.
The Three Gorges Dam concrete was made of special raw materials and mix proportions that were significantly different from those of other dams. Due to the particularity and importance of the Three Gorges Dam concrete, its mechanical properties, volume deformation, durability [4], [58], [59], [60], and AAR have been investigated by Chinese scholars over the past 20 years [61], [62], [63], [64]. However, there is still a lack of research on the very long-term (more than one year) durability and microstructure of the Three Gorges dam concrete to understand the “real” development of performances and the actual state of the microstructure. In this study, the ten-year evolution of the mechanical properties and durability of Three Gorges Dam concrete were investigated in a local area. Moreover, the impact of the service environment on the long-term development of the concrete performance was compared and investigated. Simultaneously, the microstructural characteristics of the dam concrete and the corresponding concrete drilled from the Three Gorges Dam were further analyzed using thermogravimetric analyses, X-ray diffraction (XRD), mercury intrusion porosimetry (MIP), nitrogen adsorption/desorption (NAD), and electron microscopy to understand the microstructural states of the dam concrete in the actual structure of the Three Gorges Dam.
2. Background information of the concrete in the Three Gorges Dam
2.1. Service environment of the dam concrete
Construction of the TGP began in the 1990s. The dam is located in Yichang City, in the middle reaches of the Yangtze River, China. Meteorological data in the Three Gorges Dam area shows that the daily maximum, daily minimum, and annual average temperatures from 2003 to 2006 were 39.8, −1.9, and approximately 17.0 °C, respectively. The monthly atmospheric temperature in the Three Gorges region is 5–30 °C and the water temperature is 10–26 °C [65]. The Yangtze River is a typical carbonate river, and the concentrations of the main elements in the water are listed in Table 1 [66], [67].
The environmental characteristics in the middle reaches of the Yangtze River, where the Three Gorges Dam is located, can be generally summarized as: ① The mild climate results in a small annual temperature difference; therefore, there is little risk of freeze–thaw cycles; ② because there is no particularly large gravitational potential energy difference, the water is relatively flat, in spite of the large water flow; and ③ the concentration of various ions, especially magnesium and sulfate, in the water is not high. Therefore, the effects of multiple chemical attacks on the dam concrete are not violent. In such a service environment, concrete in practical projects is mainly subjected to leaching, current scour, slight freezing damage, and potential chemical attack from trace ions over a long period, resulting in microstructural degradation and durability deterioration [43].
2.2. Concrete in different positions of the dam
The profile structure of the Three Gorges Dam can be simplified as shown in Fig. 1. Six primary engineering parts with different functions constitute the basic functions of the dam. Zone 1 is a concrete area within the dam as well as the main spatial composition of the dam; therefore, the hydration heat release process should be strictly controlled in this zone. Zone 2 is the external concrete area of the dam, which isolates the interior of the dam from the surrounding environment, and therefore interacts most directly with ambient gas and liquid. Zone 3 is part of the external concrete area but is in the area where the water level changes. Therefore, the concrete material in Zone 3 is subjected to severe dry–wet cycling, and the environmental conditions change periodically. Zone 4 is the foundation and structural concrete area. The concrete in this area is supposed to undertake the main structural mechanical function of the dam; therefore, it must have certain physical and mechanical properties. Zone 5 is part of the complex structural area, which has high requirements for the filling property of the concrete mixture; therefore, the concrete in this zone has a higher fluidity than the concrete in other parts. Zone 6 is an anti-scouring area. Concrete in this area should resist the attack and abrasion of the dam structure owing to the high water flow rate, such as flood discharge; therefore, the concrete in this zone has extremely high mechanical performance requirements. The concrete used in the engineering structure should be designed to satisfy the functional requirements of the different zones [4].
3. Experimental procedures
3.1. Raw materials
The concrete used in the construction of the TGP was strictly controlled. MHC was adopted in the dam engineering design. The cement clinker was produced by the Gezhouba Cement Plant in accordance with the provisions of Chinese national standards [28], and the 28 days compressive strength was 63.6 MPa. To compensate for the thermal shrinkage inside the dam structure, the MgO content of the clinker was controlled at approximately 3.5%–5.0% [30]. A large amount of high-quality FA was also used in the preparation of concrete materials [29]. The mass ratio of FA in the cementitious materials in the designed mix proportions was 20%–40%. The compositions of the clinker and FA are listed in Table 2 [28], [29], [30], and the particle size distribution of the binders is shown in Fig. 2.
Artificial granite aggregates from excavations or nearby quarries were crushed, processed, and screened, and used as aggregates [61], [62]. The granite was mainly composed of feldspar (60%–70%), quartz (20%–40%), and some mica (approximately 10%). Table 3 [61], [62] lists the chemical compositions of the granite aggregates. It is important to note that crushed plagioclase granite was used as the coarse aggregate and porphyritic granite was used as the artificial fine aggregate [63], [64]. Owing to its mineralogical composition, granite has potential risks of long-term alkali–silicon reactions; therefore, the alkali content of the raw materials was controlled. Table 4 summarizes the physical characteristics of the fine and coarse aggregates. The mix proportion used multi-gradation aggregates because of the different functions of concrete.
3.2. Concrete specimens
Fig. 3 shows that during the construction of the Three Gorges Dam, six types of concrete specimens corresponding to concrete in different parts of the dam [31] were prepared in the laboratory using the same raw materials and mix proportions. From the perspective of the concrete materials, the laboratory concrete specimens were almost identical to the concrete used in the actual project. Table 5 lists the mix proportions of the laboratory dam concrete. Because the high volumetric content of aggregates with a maximum size of 150 mm made the standard specimens unsuitable for testing, aggregates larger than 40 mm in the fresh concrete mixtures were removed before being cast into the mold. Furthermore, two curing conditions were adopted for the laboratory concrete, namely, standard and natural curing. It should be noted that standard curing refers to specimens that were placed in a standard curing chamber (temperature = 20 °C and relative humidity ≥ 90%) for curing to specified ages, and natural curing refers to the specimens that were moved to outdoor conditions after 28 days of standard curing for preservation in the natural environment, as shown in Fig. 4(a) and (b). It is worth mentioning that the location of the laboratory was very close to the Three Gorges Reservoir region; therefore, the natural curing conditions of the laboratory concrete were almost identical to the service environment at the Three Gorges Dam location.
The six types of laboratory concretes were used to study the effects of the two different curing conditions on the continuous long-term mechanical properties and durability of the dam concrete. The tests on the long-term properties of the concrete lasted for ten years, beginning in 2003. After ten years of property tests, the laboratory concrete was preserved. The macro-properties and microstructure of the remaining 17-year-old laboratory concrete cured under natural conditions were analyzed to study the microstructural development.
Simultaneously, concrete cores were drilled from various parts of the Three Gorges Dam for microstructural analysis to study the microstructure of dam concrete in a real service environment. As previously mentioned, these concrete cores were comprised of the same raw materials and mix proportions as the laboratory concrete. Table 6 lists crucial information regarding these cores.
3.3. Methodology
3.3.1. Continuous long-term performances of laboratory dam concrete
The long-term performance tests for the compressive strength, flexural strength, natural vibration frequency, freezing–thawing resistance, water penetration resistance, and natural carbonation were based on Chinese standard SL 352–2006 [68]. 150 mm × 150 mm × 150 mm cube specimens were used for the compressive strength test, and 100 mm × 100 mm × 400 mm prism specimens were used for the flexural strength, natural vibration frequency, freezing–thawing resistance, and natural carbonation tests.
(1) Compressive and flexural strength. The compressive and flexural strengths were tested using three cubic and two prismatic specimens at the rate of 0.4 MPa·s−1 and 110 N·s−1, respectively. The specimens were cured under standard conditions for 28, 90, and 180 days; 1, 3, 5, and 10 years. Another batch of parallel specimens was tested under natural conditions using the same experimental procedure. A four-point loading method for the simply supported beams was adopted in the flexural strength test.
(2) Natural vibration frequency. The natural vibration frequency (f) is an inherent property of solid materials. When the applied mechanical vibration frequency is equal to the natural vibration frequency of the specimen, resonance is generated, and the natural vibration frequency of the specimen is measured accordingly (Fig. 5).
(3) Resistance to freezing and thawing. Three prismatic specimens were used for the freeze–thaw cycling tests. A freeze–thaw cycle included a 2.5 h cooling process and a 1.5 h heating process. At the end of cooling and heating, the central temperature of the specimen should be controlled at (−17 ± 2) and (8 ± 2) °C, respectively. The relative dynamic moduli of elasticity (En) and mass loss (Wn) were calculated using Eqs. 1, 2.
where f0 and fn are the natural vibration frequencies before and after the manifold freeze–thaw cycling, respectively; and and are the masses of the concrete before and after the manifold freeze–thaw cycling, respectively.
(4) Resistance to water penetration. A truncated cone specimen with an upper diameter of 175 mm, a lower diameter of 185 mm, and a height of 150 mm was used to measure the water permeability resistance of the concrete. After standard curing for 48 h, the specimens were demolded and cured to specific ages. Before testing, the sides of the specimens were completely sealed in paraffin. The specimen was maintained at a water pressure of 1 MPa for 24 h during the testing process and then split to measure the water penetration height. The relative permeability coefficient (Kr) of the concrete was calculated using Eq. 3.
where Kr is the relative permeability coefficient (cm·h−1), a is the water absorption of concrete (0.03), Dm is the water-permeated height (cm), t is the water pressure maintenance time (h), and H is the water pressure in terms of the height of the water column (1 MPa = 10 240 cm).
(5) Natural carbonation. Specimens cured under natural conditions were used to measure the natural carbonation depth at different aging durations. The carbonation depth test was performed immediately after the bending strength test. A 1% phenolphthalein ethanol solution was sprayed onto the fracture surface, and the average carbonation depth at multiple points on both sides was measured after 30 s. The carbonation rate coefficient is given by Eq. 4 [69], [70].
where is the carbonation rate coefficient (mm·a−0.5), is the carbonation depth (cm), and tc is the carbonation time (year).
3.3.2. Macro-properties and microstructure of 17-year-old laboratory concrete
Regarding to macro-properties of 17-year-old laboratory concrete. The transport properties, which are important for the durability of concrete, are reflected in the rapid chloride penetration test (RCPT) and total porosity. The air void structure is a crucial parameter that affects the frost resistance of hardened concrete.
(1) RCPT. The laboratory concrete was cut into flat 100 mm × 100 mm × 50 mm (length × width × thickness) cubic specimens and were vacuum-filled with water for 24 h before the test. During the test, a 60 V direct current (DC) was applied to both sides of the axial direction of the test specimen. A 0.3 mol·L−1 NaOH solution and 0.3% NaCl solution were added to the testing cells at the positive and negative poles of the specimens, respectively. The total charge of the concrete within 6 h was recorded as an indicator of its chloride penetration resistance, as shown in Fig. 6.
(2) Total porosity. The porosity was determined from measurements of the saturated and dry weights of the concrete [71]. Three blocks of laboratory concrete in each group were first filled with water under vacuum for 24 h. The mass of the water-filled concrete was then determined and marked as . A hydrostatic balance was used to measure the mass of the concrete specimen in water and marked as . Finally, the concrete specimen was placed in an oven at (105 ± 5) °C for seven days to remove the water in the concrete pores, and the mass of the dried concrete (marked as ) was measured at room temperature. The porosity of concrete () can be calculated using Eq. 5.
(3) Air-void structure. The pore structure of the concrete was measured using the air void parameters [72]. Three 100 mm × 100 mm × 20 mm concrete specimens were cut along the axial direction from the top of the concrete specimen. The observation surface was pre-ground using 400-, 800-, and 1000-mesh SiC sandpaper. An ultrasonic cleaning instrument was used to clean the surfaces of the specimens at each step. Subsequently, the concrete surface was continuously polished for 2.5 h using an alumina powder polishing solution (powder:water = 1:3). The polished concrete specimens were placed in an oven at (105 ± 5) °C for 24 h. After the specimens were cooled to room temperature, the observation surface of each specimen was coated with a thin layer of carbon ink, and the air voids were filled with ultrafine calcium carbonate powder to enhance contrast. An NELD-BS630 fully automatic bubble parameter equipment (NELD, China) was used for the analysis. The cross-sectional area of each concrete specimen was approximately 80 mm × 80 mm, and the total length of the straight traverse wire was 6480 mm. A stepping motor was used for step-by-step image acquisition, with a step length of 1 mm and a total testing time of approximately 1 h. The minimum size of the air voids as determined using a microscope was 5 μm.
(4) MIP. In this test, mercury is forced into porous materials under external pressure. Based on the hypothesis of cylindrical pores in Washburn’s theory, a certain pore diameter is related to a certain pressure value, and the intrusion volume of mercury is equivalent to the pore volume of the corresponding diameter. Owing to the significant existence of closed pores and “ink-bottle” pores, pore structure errors are inevitable in MIP tests for cementitious materials [71], [73]. The MIP was measured in this study using an MK-AutoPore IV 9510 porosimeter (Micromeritics, USA). The maximum and minimum pressures of the mercury intrusion porosimeter are 400 MPa and 1.4 kPa, respectively, corresponding to a pore diameter range of 3 nm–400 μm. Six grams of 3–4 mm crushed concrete blocks were used for the MIP test, and as many coarse aggregates as possible were removed during the sampling process. The samples were immersed in isopropanol for three days and then dried at 45 °C for five days in a vacuum before the MIP test.
(5) NAD. The samples for the NAD tests were cut into small 1–2 mm diameter pieces. The samples were immersed in isopropanol for three days and then dried at 45 °C for five days in a vacuum before the NAD test. The micropore size distribution of the concrete specimens was determined using a TriStar II 3020 device (Micromeritics, USA) at 77 K and calculated using the Barret–Joyner–Halenda (BJH) model according to the N2 desorption curves. It should be noted that the samples were degassed under vacuum at 100 °C for 6 h in this test.
(6) Mineralogical analysis. The concrete samples were carefully broken into blocks, and distinguishable aggregates were removed as far as possible. The residual mortar samples were ground into powders of appropriate fineness using a freeze-grinding mill. The mineralogical compositions of the concrete were analyzed using XRD. A Bruker D8-Advance X-ray diffractometer (Bruker, USA) was used to perform the analysis at a scanning speed of 5°·min−1 in the scanning range of 5°–70°. Additional mineralogical variations in the concrete were determined using thermogravimetric analysis (TGA). Powder samples with a mass of approximately 15 mg were heated from room temperature to 1000 °C at 10 °C·min−1 in a nitrogen atmosphere. The mass-loss curve was plotted as a function of the temperature, and derivative thermogravimetric (DTG) curves were obtained.
(7) Scanning electron microscopy (SEM) analysis. Approximately 3 mm thick concrete slices were immersed in isopropanol for seven days to remove the water in the pores, and then were dried to a constant weight in a vacuum drying oven at 45 °C. The dry samples were impregnated with a low-viscosity epoxy resin in vacuum. Because the depth of the impregnated epoxy was only a few hundred microns, the sample grinding time should not be lengthy to avoid excessive grinding. The samples were ground using 300-, 600-, and 1200-grit SiC sandpaper for 5 min each. The samples were then polished using successive diamond suspensions with particle sizes of 9, 3, 1, and 0.25 μm under a force of 40 N [74]. Each polishing step respectively lasted 30 min, 60 min, 2 h, and 4 h. At the end of each polishing step, the samples were ultrasonically cleaned with ethanol to remove any polishing residue. Kerosene was used to cool the samples during the polishing process. The polished samples are shown in Fig. 7. Each polished sample was coated with gold before SEM observation using a vacuum coating machine. An FEI (USA) Quanta-200 SEM equipped with energy dispersive X-ray spectrometry (EDS) was used to conduct scanning electron microanalyses. The electron-beam acceleration voltage was 15 kV, and the beam spot size was 5.0 μm in the backscattered electron (BSE) image acquisition and 6.0 μm in the EDS and element mapping analysis. The magnification of the BSE images in the element mapping analysis was set as 1000 × and the size of these images was 1024 pixels × 943 pixels (each pixel size is 0.291 μm).
3.3.3. Microstructure of the dam concrete cores
The dam concrete cores were microscopically analyzed to compare the microstructural differences between concrete in actual service environments and naturally-cured laboratory concrete. The characterization methods for the microstructure of the dam concrete cores were the same as those for the 17-year-old laboratory concrete. XRD, thermogravimetry (TG)/DTG, MIP, NAD, and SEM were also used to characterize the microstructures of the concrete cores.
4. Results and discussion
4.1. Mechanical properties and durability of laboratory concrete
4.1.1. Compressive and flexural strength
Fig. 8 shows the compressive strengths of concrete with different aging durations under the two curing conditions. The compressive strength of all concretes increased with age. The compressive strength increased rapidly in the first year, as compared with the strength one year later. After one year, the development of the concrete strength entered a relatively slow period, which is due to the exhaustion of highly reactive minerals in the cement clinker, such as C3S and C3A, as well as the hydration of partial C2S and ferrite [75]. The lack of space available for hydrate formation also results in the deceleration of strength growth [73]. In the case of the standard-cured concrete, the strength of concrete blended with no less than 30% FA continued to increase after one year; however, the compressive strength of concrete blended with < 30% FA was stable. This phenomenon was primarily attributed to the chemical effects of FA [76], [77]. Comparing the effects of the curing conditions on strength, the strength of naturally-cured concrete was slightly higher than that of standard-cured concrete within the first year, whereas the strength of the standard-cured concrete exceeded that of the naturally-cured concrete after one year. The pores were filled with carbonates caused by carbonation, which might have resulted in a further increase in the compressive strength during the first year [33]. However, it was more significant that standard curing promoted an increase in concrete strength over a fairly permanent period because the loss of portlandite, a reactant produced by natural curing, limits the long-term reaction of FA [78]. Furthermore, the formation of microcracks owing to weathering weakens the contribution of FA to the strength.
The water–binder (w/b) ratio and FA content are crucial factors for the strength and strength growth of concrete. Concrete with a lower w/b ratio has a higher initial strength. The strength of concrete with a w/b ratio of 0.35 and 20% FA content at 28 days and 10 years were significantly higher than those of the other concretes. However, throughout the strength development period, the strength growth was significantly lower than that of the concrete blended with a higher FA content. This suggests that an increase in FA content enhances the range of strength growth of concrete during its entire service life because of the continuous pozzolanic reaction [76].
Fig. 9 shows the flexural strength of concrete with different aging durations under the two different curing methods. After ten years of standard curing, the flexural strength tended to stabilize. The initial flexural strength of the concrete increased with a decreasing w/b ratio. However, the ultimate flexural strength values of all concrete specimens were stable at approximately 7 MPa. Concrete with a higher w/b ratio and FA content exhibited a greater increase in the flexural strength because a higher initial porosity provided more filling space for the FA hydration products [73]. Unlike the compressive strength, there was no significant difference in the development of the flexural strength between the naturally- and standard-cured concretes.
4.1.2. Natural vibration frequency
The natural vibration frequency is an inherent property of materials. For concrete, which is a porous solid material, the natural vibration frequency reflects its intrinsic mechanical properties. The dynamic elastic modulus can be calculated using the natural vibration frequency [79]. Fig. 10 shows the natural vibration frequencies of concrete with different aging durations. The natural vibration frequency (28 days) was related to the w/b ratio and FA content. The initial natural frequency (28 days) clearly tended to increase with a decreasing w/b ratio, except for the high-fluidity concrete “0.45FA20*-L.” Although the w/b ratio and FA content were the same, the high-fluidity concrete contained more pastes and a smaller volume fraction of aggregates. Aggregates generally possess a higher hardness and modulus of elasticity; thus, concrete with a lower aggregate content has a lower natural frequency [4]. Under standard curing conditions, the natural frequency increased rapidly in the first stage, and then the rate of increase markedly slowed after one year, which was consistent with the compressive strength. The natural vibration frequency of all the concretes was more similar to the prolongation of age owing to a higher increase in the natural frequency of concrete blended with more FA. The natural vibration frequency of the naturally-cured concrete increased slightly during the first 90 days; hence, the natural frequency of the naturally-cured concrete was clearly lower than that of the standard-cured concrete during the entire period. However, the reduction in the natural vibration frequency caused by natural curing was < 5% for all the concretes.
4.1.3. Resistance to freezing and thawing
Fig. 11 shows the dynamic elastic modulus loss and mass loss of concrete after several freeze–thaw cycles. Table 7 lists the number of freeze–thaw cycles at different aging durations. All concretes exhibited a significant difference in the mass loss after freeze–thaw cycling; therefore, it is more appropriate to evaluate the frost resistance using the relative dynamic elastic modulus. Under standard curing conditions, frost resistance was not significantly affected during the curing period. The loss of the dynamic elastic modulus of concrete with a low w/b ratio did not exceed 20%, whereas the loss of the dynamic elastic modulus of concrete with a w/b ratio > 0.50 was up to 30%–40%. Although numerous air voids were introduced into fresh concrete (Table 5) to enhance its freezing resistance [72], the increase in capillary pores due to the increase in the w/b ratio intensified the damage caused by ice crystallization [4]. The impact of FA on the freezing resistance was not significant for the standard-cured concrete.
The freezing resistance of the naturally-cured concrete was significantly reduced. The naturally-cured concrete only required dozens of cycles to exhibit significant losses in both the elastic modulus and mass. Owing to the accumulation of fatigue damage caused by changes in the ambient temperature and humidity, the resistance to freezing of concrete decreases over time, which can been explained by additional micro freezing–thawing effects that are produced by the drying–wetting cycles in the natural environment, causing the critical pore diameter to vary [40]. The freezing resistance of concrete with a low w/b ratio, as compared to that of the other concretes, decreased significantly after natural curing, which seemingly indicates that the frost resistance of concrete with a higher strength is more sensitive to the effects of the environment. FA may play an important role in compensating for the loss of freezing resistance [4], [80].
The mass loss of the standard-cured concrete decreased slightly with an increasing curing age, which was not consistent with the variation in the dynamic modulus of elasticity. It appears that both the w/b ratio and FA content affect the mass loss of concrete. The relative dynamic modulus of elasticity of the naturally-cured concrete decreased to 60% or lower after 50 freeze–thaw cycles; however, the mass loss was not significant. The uncertainty and complexity of the mass variation can be partially attributed to the intake of CO2 by carbonation [32], [33], [36]. Thus, mass loss cannot accurately reflect the internal damage to the concrete due to the freezing–thawing process.
The frost resistance test results of the concretes with different curing durations over five years show that the curing method is important for the frost resistance of concrete. The strength of ten-year-old concrete after freeze–thaw cycling was measured to understand the relationship between the relative dynamic elastic modulus and strength loss (Table 8). Fig. 12 shows the relationship between the dynamic elastic modulus loss and strength loss after freeze–thaw cycling. The strength loss of concrete is a function of the dynamic elastic modulus loss. Although the dynamic elastic modulus of concrete decreased rapidly with an increasing number of cycles, the strength losses were limited to a concentrated scope. The average losses in compressive strength and flexural strength of the naturally-cured concrete were 26% and 42%, respectively. Moreover, the flexural strength loss of the naturally-cured concrete was higher than that of the standard-cured concrete, whereas the compressive strength losses of both concretes showed little difference. This indicates that freeze–thaw cycling increases the brittleness of concrete [4]. Additionally, the number of freeze–thaw cycles for the naturally-cured concrete was reduced by 91%, as compared to the standard-cured concrete, and the grade of frost resistance decreased rapidly.
The relative dynamic elastic modulus of ten-year-old concrete after freeze–thaw cycling was 66.24%–82.54%, and the residual strength of the naturally-cured concrete was also higher than the designed strength when the number of freeze–thaw cycles was 15–80. These results indicate that concrete damaged by freeze–thaw cycling is still adequate for structural bearing [4], [31]; however, the shear failure of structures will require attention because of the increase in the brittleness of the concrete.
4.1.4. Resistance to water penetration
In this experiment, concretes corresponding to three parts of the dam were adopted as experimental subjects. These concretes emphasize the resistance of water permeability. Fig. 13 shows the water penetration height and coefficient of concrete over one year. The penetration coefficients of the standard-cured concretes were approximately 2.7 × 10−7 cm·h−1, revealing their outstanding resistance to water penetration. It should be noted that anomalous test values were obtained for concrete with a w/b ratio of 0.35. These results differ from those of other studies [81]. However, other researchers reported that the permeability coefficient was significantly affected by the w/b ratio only when the w/b ratio was > 0.6 [4]. Therefore, the water penetration heights and coefficients in this study exhibited no obvious relationship with the w/b ratio. Additionally, the impacts of FA, aggregate content, air-entraining agent, and sample preparation are also significant for the water penetration resistance of concrete [4]. Moreover, the water penetrability of the naturally-cured concrete increased to different degrees, as compared to the standard-cured concretes; nevertheless, both concretes had nearly the same strength (Fig. 8). This indicates that strength is not directly related to the penetrability of concrete because of the difference between connected and closed pores [4], [71]. Natural curing resulted in more connected microcracks in the concrete with higher porosity, resulting in the phenomenon where the penetrability increase caused by natural curing decreases with a decreasing w/b ratio. Although the concrete was impermeable after early standard curing, the continuous attack under natural conditions continued to exert adverse effects on the concrete.
4.1.5. Natural carbonation
Fig. 14 shows the carbonation depth of the concrete under natural conditions. As the natural curing time increased, the carbonation depths in all groups of concrete increased continuously. The ten years carbonation depths of all concretes did not exceed 8 mm, which is significantly less than the carbonation depth of concrete after accelerated carbonation in the laboratory [32], [33], indicating that the carbonation rate of concrete in a real natural environment is relatively slow. The carbonation depth increased with an increasing w/b ratio, which was mainly due to the high porosity caused by the high w/b ratio [33], [82]. Furthermore, a sudden increase in the carbonation depth for “0.55FA40-L” was observed over one year. This may be due to accelerated carbonation caused by the inevitable cracking of the specimen [4].
Fig. 15 shows the carbonation rate coefficient results. The carbonation rate at 28 days was significantly higher than that at one and ten years, revealing that the diffusion of CO2 in the concrete gradually slowed with an increasing carbonation depth [34]. The average carbonation rate decreased with a reduction in the w/b ratio, particularly for concrete aged for less than one year. In another study [69], concrete blended with FA exhibited a reduced carbonation resistance. However, the data in this study indicates that the influence of the FA content on the carbonation of concrete was not significant. Moreover, the permeability resistance of concrete may be affected by variations in its microstructure caused by natural carbonation [32], [83].
4.2. Macro-properties of the 17-year-old laboratory concrete
4.2.1. Rapid chloride penetration
Fig. 16 shows that the total charge of all the specimens were within the range of 100–300 C after 17 years, regardless of the curing mode, w/b ratio, and FA content. According to the American Society of Testing Materials (ASTM) C1202 classification [84], the permeability of chloride ions in the 17-year-old concrete was extremely low. The classification also indicated that the 17-year-old concretes had good water permeability resistance [85]. When the FA content did not exceed 20%, the porosity of the concrete decreased with a decreasing w/b ratio [71], [83], hindering the diffusion of chloride ions into the concrete. When the w/b ratio was fixed at 0.45, an increase in the FA content optimized the pore structure of the concrete, further optimized the interfacial transition zone (ITZ), and weakened the diffusion transmission paths of the attack medium [86], [87], resulting in a decrease in the passing charge. In particular, there was a particularly large increase in the electrical flux when the FA content decreased from 30% to 20%, and the passing charge of the concrete with 20% FA was generally higher than that of the concrete with 30% FA. This suggests that if the FA dosage increases from 20% to 30%, its contribution to chloride diffusion resistance is significantly improved. The effect of the w/b ratio was compensated by the contribution of FA. The RCPT results of the concretes generally revealed that the diffusion of chloride ions in the concrete was significantly affected by the pore structure [76], [87], whereas a high FA content in the concrete enhanced the effective binding of chloride ions [88]. Numerous air voids were introduced into the high-fluidity concrete, significantly increasing the porosity of hardened concrete. Additionally, an improvement in the sand ratio increased the specific surface area of the aggregate and the total volume of the ITZ. Furthermore, the large amount of cementing material used in pumped concrete increases the likelihood of cracks, which, in turn, increases the number of chloride ion transport paths [4]. Thus, easier chloride diffusion resulted in an increase in the passing charge of the pumped concrete, as shown for 0.45FA20* in Fig. 16.
4.2.2. Total porosity
Fig. 17 shows the porosity of concrete aged for 17 years. The total porosities of all concrete samples were 5.8%–8.0%. The porosity of the concrete decreased from 7.6% to 6.0% when the w/b ratio decreased from 0.55 to 0.45; however, there was only a slight variation in porosity even though the w/b ratio continued to decrease. This indicates that, in the case of a large w/b ratio, reducing the w/b ratio is more effective for the densification of concrete. The porosity of concrete is also affected by the FA content [71], [73]. The results of the total porosity can explain the mechanical properties and anti-permeability performance of the naturally-cured concrete mentioned above [89], but the frost resistance of concrete cannot be attributed simply to variations in porosity [40]. The pumped concrete had more capillary pores per unit volume because of its larger paste volume. Comparing the results of the RCPT, the reduction in porosity was an important reason for the decline in the passing charge of the concrete. However, there was little difference in the total porosity between “0.45FA30” and “0.45FA20,” which is not enough to explain the prominent increase in the passing charge. Therefore, it can be asserted that increasing the FA content from 20% to 30% significantly increases the chloride penetration resistance of concrete [86]. Hence, 30% is a very important FA content to optimize concrete durability, at least for the chloride resistance of concrete.
4.2.3. Air-void structure
Fig. 18 shows the air-void factors for various types of concrete. The air void spacing factor is strongly correlated with the freezing resistance of concrete [72]. The air void spacing factor of the 17-year-old concrete was 100–220 μm, indicating that the naturally-cured concrete retained a relatively good frost resistance. Both the w/b ratio and FA content affect the air void spacing factor to varying degrees by influencing the capillary porosity [71], [90]. Under natural curing conditions, the air void spacing factor exhibited an overall decreasing trend with an increasing w/b ratio and FA content. When the w/b ratio was increased from 0.50 to 0.55 and the FA content was increased from 30% to 40%, the air void spacing factor exhibited the maximum reduction. The pumped concrete, as compared with normal concrete, had a smaller air void spacing factor; therefore, the pumped concrete had better frost resistance (Fig. 11 and Table 8). However, the strength of the concrete specimens with a high w/b ratio was relatively lower than that of the other concrete specimens, and the residual strength and relative dynamic elastic modulus of concrete after freeze–thaw cycling were consequently poor, even with smaller air void spacing factor values [4]. Interestingly, continuous standard curing can reduce the air void spacing in concrete. The reason for this phenomenon is discussed later.
Fig. 19(a) shows the air void content in the hardened concrete. The variation in the air void content was similar to that of the total porosity. The air void content first decreased with a decreasing w/b ratio and FA content and then remained stable once the w/b ratio decreased to 0.45. The air void content of the hardened pumped concrete was slightly higher because of the greater amount of paste. It should be noted that standard curing slightly reduced the air void content, but the air void spacing factor decreased as well, which may be caused by changes in the air void characteristics. Additionally, concrete loses part of its air void content during hardening. Fig. 19(b) shows the air void loss in the concrete. When the w/b ratio and FA content were both relatively high (w/b ≥ 0.45 and FA ≥ 30%), the air void loss significantly increased with a decreasing w/b ratio and FA content. However, when the w/b ratio and FA content were further reduced, there was no significant change in the air content loss in the concrete.
The air content is not the most direct factor affecting the frost resistance of concrete; however, the morphology and distribution of air voids in concrete are [4], [72]. Hence, the total number and average radius of the air voids were measured, and the results are shown in Fig. 20(a) and (b), respectively. The trend of the number of voids was very similar to the air void content. The air content and number of voids in the hardened concrete continuously decreased with a decreasing w/b ratio and FA content, and this trend was prominent when the w/b ratio and FA content were > 0.45 and ≥ 30%, respectively. This indicates that increasing the w/b ratio and adding FA can introduce more air voids into the concrete, which is an important reason for the change in the air void spacing factor. Additionally, the pumped concrete in the plastic state has a higher air content; therefore, there are more bubbles inside the concrete after hardening, resulting in a higher total air content. Although the number of air voids increased under standard curing conditions, the total air content decreased (Fig. 19). Fig. 20(b) shows that there were no significant variations in the average radius of the air voids in all concretes, except for “0.50FA30,” and the average radius of the air voids for all concretes was approximately 100 μm. A comparison of the pumped and ordinary concretes with the same w/b ratio and FA content revealed that the difference in bubble size was not significant. More air voids were introduced in the pumped concrete; however, it had little influence on the size of the air voids. Furthermore, the size of the air voids clearly influenced the curing mode. Standard curing not only increased the total number of voids in hardened concrete but also reduced the diameter of the air voids. This led to a decrease in the air void spacing factor of concrete without an increase in the total air content. Additionally, refinement of the air void structure reduces the potential damage of coarse pores on the mechanical properties [71].
4.3. Microstructural analysis of 17-year-old laboratory concrete and dam concrete cores
4.3.1. Micropore structure
The cumulative pore distribution and differential pore distribution in two different pore diameter ranges for the 17-year-old laboratory concrete and dam concrete cores were measured using the MIP and NAD methods, respectively. The entire pore distribution for each sample was obtained by the superposition of the results from both testing techniques. For all concrete samples, pores with a diameter of 1.7–10.0 nm were characterized using NAD, and MIP was used to characterize the pores with a diameter > 10 nm. It should be noted that pores > 100 μm were considered to be visible air voids and were not counted as concrete micropores (Section 4.2.3). Based on the study by Zeng et al. [71], the pore size distribution was divided into different intervals according to diameter, that is, gel pores (up to 5 nm), mesopores (5–50 nm), middle capillary pores (50–100 nm), large capillary pores (100 nm–10 μm), and smaller air voids (> 10 μm).
Fig. 21(a) and (b) show the cumulative porosity and pore-size distribution curves, respectively, of the 17-year-old naturally-cured concrete. In the micropore range, naturally-cured concrete demonstrated a similar porosity of 6%–7%, which was significantly lower than that of the 90–360 days OPC or LHC pastes blended with FA (w/b ratio of 0.30–0.50) [71], [73], [91]. This indicates that the dam concrete exhibited a dense microstructure even after 17 years of exposure to the local environment. For pores < 100 μm, the effect of the w/b ratio on the porosity of the 17-year-old concrete was not clear. However, the total porosity obtained using gravimetry still showed a dependence on the w/b ratio, as mentioned above. It should be noted that microporosity (capillary pores) is seemingly a function of the FA content, which indicates that FA had a significant impact on the micropore structure of long-term concrete [92]. Focusing on the pore-size distribution curves of concrete cured in open air, a complex distribution of peaks corresponding to pores of various diameters is shown in Fig. 21(b). There was no obvious critical diameter for any of the curves, as described in the literature [73], [74]. The broad continuous peaks in the distribution curves of the pore entry diameter indicate the ubiquity of unconnected capillary pores of different sizes inside the concrete. Hence, the pore volume curves exhibited a steady increase with a decreasing pore size. In contrast to the paste, porous ITZs between the aggregates and the paste provided extra pores, resulting in increased porosity and additional peaks in the distribution of pore diameter (DPD) [93]. The precipitation of carbonates owing to mild carbonation has minor effects on capillary pores [32], [78], [82]. The distribution of the gel pores was also relatively homogeneous and exhibited opposite trends to those of the coarse capillary pores. Similarly, the distribution and content of gel pores did not seem to have a significant relationship with the initial w/b ratio; however, they were related to the FA content. The reaction of more FA led to the formation of more calcium (alumino)silicate hydrate (C-(A)-S-H) gels, which promoted an increase in gel pores and refined coarse capillary pores [77], [94]. Fig. 21(c) and (d) show the cumulative porosity and pore size distribution curves of the core samples, respectively. The core samples, as compared to the naturally-cured concrete, exhibited a slight increase in porosity. A total porosity of 8%–12% was observed in the core samples. The service environment of concrete is more complicated in actual dams than under natural exposure conditions. Leaching and carbonation occur continuously during the service of dam concrete, causing variations in its microstructure [56]. The exterior of the dam structure, particularly for portions permanently below the waterline, and the ordinal dissolution of hydrates increase the porosity of the concrete in this region [43], [57]. The core sample “0.50FA30” obtained from the exterior of the dam possibly had been exposed to long-term water pressure, resulting in increased porosity. The external concrete of the dam, as compared to the naturally-cured concrete, exhibited a frequent pore size distribution within 30–200 nm, as shown in Fig. 21(d). This can be explained by the calcium hydroxide firstly dissolved during the leaching of concrete, which causes an increase in capillary pores in the microstructure [37]. The gel pore content presented an opposite trend to that of the coarse capillary pores in the cores. Prominent gel peaks with a diameter of approximately 4 nm were visible in several curves, suggesting a stable C-(A)-S-H gel structure in actual concrete [71]. However, alteration processes, such as leaching and carbonation, that cause coarsening of the capillary pores also affect the gel structure [32], [37]. This resulted in a synergistic variation in the gel pore distribution with the capillary pore distribution, as shown in Fig. 21(d).
Briefly, concrete mortar cured in an extra ventricular environment is characterized by low porosity and a relatively broad pore size distribution. The w/b ratio seemingly had no impact on the micropore structure of the long-aging duration concrete, but the initial FA content did. The micropore structure of the concrete sampled from multiple parts of the dam indicates that the actual engineering environment led to a significant degradation of the microstructure. In this study, the porosity and capillary pore volume of the core samples increased to varying degrees, as compared to concrete preserved outdoors. A sample of the exterior concrete revealed an increase in porosity, which was attributed to the continual leaching process. This indicates that microstructural degradation caused by leaching may be a severe problem for the concrete in the Three Gorges Dam.
4.3.2. Mineralogical composition
(1) TGA and DTG.Fig. 22 shows the TGA and DTG curves of the powders from the laboratory concrete and dam concrete cores. The complicated mass loss of concrete at 50–1000 °C originates from dehydration or/and decarbonation [74]. Within this temperature range, the mass loss of concrete was 5%–10%, which is lower than that of the pastes because of the presence of aggregates [92], [95]. All laboratory concrete samples exhibited a first significant mass loss at 50–200 °C, which was mainly due to the dehydroxylation of C-(A)-S-H gels and interlayer water loss [43]. Furthermore, the decomposition peaks at approximately 100 °C, due to the water loss of ettringite [95], were not detected in the DTG curves. Furthermore, the bound water content (mass loss at 50–550 °C) clearly increased for all concretes with a decreasing w/b ratio and FA content. This can be explained by the increase in binders in the concrete (Table 5). No typical mass loss related to variations in the AFm phases, such as monosulfate (Ms) and monocarbonate, was observed at 150–300 °C. The expected dehydration peak of brucite [25] at 350–400 °C was not observed for all samples because of the complex and indefinable mass loss at 200–370 °C [74]. The second significant peaks in the DTG curves at approximately 400 °C were due to the decomposition of portlandite. This was clearly visible in the DTG curves of the laboratory concrete but was barely visible in the DTG curves of the dam concrete cores, which demonstrates that almost no portlandite existed in the dam concrete. The intensity of the portlandite decomposition peak decreased with an increasing FA content, suggesting that the adequate reaction of FA effectively consumed the portlandite produced by the hydration of cement [77], [96]. Additionally, blending more than 30% FA can fully consume portlandite to alter the phase assemblages of the composites at a later period [94], [97]. The third peak in the DTG curve, accounting for the majority of the total mass loss of concrete, occurred at 550–700 °C because of the decomposition of calcite. Calcite is mainly derived from the associated minerals of the aggregates, and some carbonates may be included because of the slight carbonation of concrete [32].
(2) XRD. Four types of aggregates that could be distinguished by the naked eye were selected from the cores and ground into powders for XRD analysis (Fig. 23). The aggregates mainly consisted of quartz, albite, clinochlore, annite, and calcite. In addition to quartz and calcite, the other phases were primarily aluminum silicate minerals. Overall, the concrete aggregates were mainly acidic granite; therefore, the alkali-silicate reaction should be prevented. Previous studies indicated that the use of these aggregates did not cause serious AAR damage in the TGP [4], [61], [62]. Fig. 24 shows the XRD patterns of the laboratory concrete and concrete cores. Narrow peaks with high intensities were observed in these patterns, indicating that the concrete had numerous well-crystallized phases. Generally, it is difficult to accurately identify anhydrous or hydrated compounds in concrete because of the presence of aggregates [43], [57]. Previous XRD analyses of the aggregates can help determine the mineral composition of concrete. Most diffraction peaks in the diffraction patterns of the concrete samples were preliminarily identified as aggregates.
Quartz was the most common phase in concrete, with the vast majority originating from the aggregates. Calcite was present in all concrete samples; however, it was almost impossible to determine whether calcite originated from the carbonation reaction and/or carbonates in the aggregates. Portlandite was present in all the laboratory concrete samples. However, it was barely detected in any of the concrete cores collected from the dam site. The intensities of the portlandite diffraction peaks in the laboratory concrete decreased with an increase in blended FA; in particular, the downward trend of the peaks at 18° was the most significant. This reveals that the depletion of portlandite increased with an increasing initial FA content, which is consistent with the tendency of the thermal analysis. FA reacts slowly and continuously with portlandite to form amorphous C-(A)-S-H gels [77], [95]. The low activity of siliceous FA prevents it from fully reacting, even after a long hydration period [76], [97]. Therefore, some portlandite remained in the concrete blended with no more than 20% FA, as shown in the DTG curves. After more than 17 years of hydration, the mineral phases in the clinker were almost exhausted. Therefore, the main phases of the clinker, such as alite, belite, and aluminate, were not detected in the patterns. Combined with the micropore structure results (Fig. 21), it was demonstrated that there was insufficient space and reactants for hydration [73], [98], indicating that the cement had almost reached its maximum degree of hydration. Significantly, the microstructure of concrete consisted of Ms instead of ettringite because of the high Al2O3 content of FA [95], [96]. Moreover, because of the relatively low magnesium content, no magnesium-rich phases, such as periclase and brucite [25], were discovered in the XRD patterns.
4.3.3. Microstructure
(1) Naturally-cured concrete. Fig. 25 shows BSE images of the 17-year-old naturally-cured concrete specimen. The observation area for each sample was selected to include as much FA, ITZs, and magnesium phases as possible.
The microstructures of all the concretes were relatively dense, which is consistent with the micropore structure data. Most of the closed pores were due to the hollow structure of FA. Almost no ITZ was observed at the edge of the aggregates wrapped in the gel. However, porous regions and ITZs were observed in the concrete because of the accumulation of numerous FA particles [99]. The clinker minerals in all concretes were almost completely hydrated to form abundant C-(A)-S-H gels. The microstructures exhibited a loss of distinction between the inner products (Ip) and outer products (Op), as described in many studies [99], [100], [101]. The uniformity of the gray values in the BSE images also represents the similarity in the chemical compositions of the C-(A)-S-H gels [74]. These results indicate the formation of highly homogeneous C-(A)-S-H in the microstructure of blended concrete at very long aging durations. Portlandite was hardly observed in the BSE images, as compared to the TG and XRD results. Fully or partially reacted FA was embedded throughout the entire microstructure, particularly in concrete blended with a high FA content. The spherical FA particles were wrapped in dense gels. Smaller FA particles were surrounded by Op layers. Obvious dark transition layers were observed between the residual FA and Op layers, indicating the existence of loose Ip [95]. However, the thickness of the Ip region was extremely nonuniform for each FA particle. Additionally, the larger FA particles only had very thin product layers, demonstrating a very low degree of reaction for larger FA. The different post-hydration states of FA depend on its heterogeneous physical form and chemical composition [102]. From the elemental mapping, the vast majority of aluminum was concentrated in the FA and its Ip [103] as well as the aggregates. Other aluminum dissolved from FA was evenly distributed in C–(A)-S-H gels and other aluminum-rich phases [77]. Many fence-like products were formed in the pockets, as observed in the BSE images, particularly in Fig. 25(a). The EDS analysis revealed that the fence-like phases were Ms, which is in agreement with previous XRD results. Because FA contains a large amount of Al2O3 but little SO3, the ettringite in the system is gradually converted to Ms [77]. Therefore, the existence of Ms in the concrete blended with FA was more easily observed than that with ettringite.
The mass fraction of MgO in concrete was relatively low [31]; therefore, it was difficult to detect periclase and brucite using XRD and TG analyses. However, magnesium could be identified using BSE images and EDS mapping. Numerous dark magnesium-rich crystals were embedded in the matrix. The MgO crystal particles were located close to ferrite and exhibited different morphologies and sizes. The majority of MgO crystals appeared as irregular polygons with a size of 5–20 μm. Additionally, the fence-like MgO shown in Fig. 25(c) was noticeable. It should be noted that bulk periclase was not fully hydrated, even after 17 years. The grey value decreases once the periclase is hydrated to form brucite [24] and can also be demonstrated by the attachment of oxygen based on mapping. The degree of reaction of periclase particles depends on their spatial positions [25]. As shown in Fig. 25(e), the periclase in the upper part of the BSE image was wrapped in an iron-rich phase and hardly reacted with it. However, the periclase in contact with the C-(A)-S-H gels was partially converted to brucite. Overall, the average degree of periclase hydration was low.
(2) Core samples of the dam under real service conditions.Fig. 26 shows BSE images and elemental maps of various concrete cores obtained from the dam. The microstructure of the core samples was not as dense as that of the laboratory concrete. Many porous areas were mixed, with a variety of phases existing between the aggregates. As mentioned above, FA rich in aluminum and silicon remained visible in the BSE images and elemental distribution maps.
It is noteworthy that the concretes from the inside of the dam exhibited a region rich in magnesium in the left half of the field of view in Fig. 26(a). The region contained magnesium, iron, and calcium, but almost no aluminum or silicon. Therefore, this zone can be considered a magnesium-rich nest where periclase particles are mixed with ferrites. However, the particle sizes of these periclase crystals were significantly smaller than those of the normal periclase particles mentioned above. Song et al. [104] also reported the presence of periclase nests in high-magnesium clinker, which were mainly caused by inadequately ground dolomite. Sulfur was also concentrated in this region, indicating that sulfates (probably gypsum) precipitated there [37]. Furthermore, Fig. 26(a) contains a dark area rich in magnesium below the field of view. The average size of this phase was larger than that of the periclase particles mentioned above; hence, it was considered to be a magnesium-rich gel rather than anhydrous or hydrated periclase [105].
Fig. 26(b) shows the microstructure of a core from the dam foundation. The void spaces in the vicinity of the aggregates were not filled with hydration products. A low-density ITZ was not observed around the tightly wrapped aggregate. The FA particles were in close contact with the C-(A)-S-H gels, and their hydration states were consistent with those discussed above. Portlandite residue was also found in the middle of the field of view, and the enrichment of calcium confirmed the presence of portlandite. Very few periclase particles are visible in this figure, and oxygen was also concentrated here, which indicates sufficient hydration of these periclases.
Fig. 26(c) shows the microstructure of a concrete core from the exterior of the dam. Multiple hydration products and FA residuals appeared to be loosely stacked between the aggregates, which resulted in the formation of a very porous microstructure. The porosity of the concrete from this zone was significantly higher than that of the concrete from other parts of the dam, which was verified by the results of the pore structure above. Previous studies [78], [106] on the effects of leaching and carbonation on pastes blended with FA reported a loose microstructure similar to that of the concrete taken from this zone in the dam. Leaching and carbonation due to exposure to river water were the main causes of the coarsened microstructures of the exterior concrete [43]. Moreover, higher porosities were observed in the zones near the aggregates, whereas the zones far from the edges of the aggregates appeared to have lower porosities. It can be clearly seen that the grey value between the aggregate and paste decreased, revealing an increase in the porosity of the ITZ. Hence, the alteration of concrete has a significant effect on the ITZ [4].
Fig. 26(d) shows the microstructure of the dam concrete near the water line. Owing to less FA in the field of view, the microstructure of the paste was very compact, as compared to that of the other core samples. Except for the individual pores, there were also enclosed zones that contained many pores inside and were tightly embedded in the C-(A)-S-H gels. Moreover, residual ferrite and iron-containing siliceous hydrogarnets were distinguishable because of their bright gray level [107]. The porosity of the paste decreased with an increasing distance from the aggregate edge. Dozens of microns of porous layers were visible at the aggregate-paste boundary, indicating the existence of weak ITZs. Approximately 10 μm periclase particles were visible in the C-(A)-S-H matrix. The outside of the periclase in contact with the C-(A)-S-H gels was converted to brucite; however, the reaction of the interior of the periclase was hindered due to the lack of access to water [25].
C-(A)-S-H gels were generally the primary product occupying the vast majority of the microstructure in the concrete cores, and FA played a secondary role in the spatial composition of the microstructure. Ms, an iron-containing phase, and infrequent portlandite were observed in the microstructure. The reacted forms of FA were similar in both the actual and laboratory concretes. However, the microstructure of the laboratory concrete was generally denser than that of concrete used in actual engineering applications. Numerous loose porous regions were visible in the microstructure of the concrete in the dam. Additionally, the compactness of the concrete microstructure in different parts of the dam varied significantly owing to the differences in the service environment.
5. Further discussion
5.1. Dense microstructure: Origin of the good properties of long-term dam concrete
The continuous development of the properties of the laboratory concrete over ten years indicated that the dam concrete had long-term service performance, including excellent mechanical properties and durability, regardless of whether it was cured in natural or standard environments. Despite the adverse effects of long-term natural curing on the frost resistance of concrete, the structural requirements were still satisfied. The properties of the 17-year-old naturally-cured laboratory concrete also confirmed this inference. It is well known that the performance of concrete depends on its microstructure [98]; therefore, good durability can be explained in terms of the microstructure of concrete.
The pore structure is an important component of the microstructure and is frequently used to reflect the link between the microstructure and its properties [32], [40], [91]. A basic understanding of the microstructure of the long-term concrete, based on the composition and multiscale pore structures of the concrete, is presented in this study. Because of the existence of ITZs, capillary pores with different diameters were widely distributed in the microstructure of the dam concrete. The features of typical pore-size distribution curves [71], [73], [74], such as the critical entry size, are covered by these additional coarse capillary pores [93]. Despite the complicated pore distribution, the microporosities of all concretes were very low. As shown in Fig. 27 [51], [54], [56], [108], [109], [110], [111], [112], [113], [114], the microporosity of the hydraulic concrete applied in the Three Gorges Dam was compared with that of other long-term concretes. The concretes for comparison were made using different raw materials, such as Portland cement, slag cement, FA cement, and other supplementary cementing materials, and were exposed to various extreme conditions for long periods [51], [54], [56], [108], [109], [110], [111], [112], [113], [114]. According to the literature, these concretes exhibited superior mechanical properties and durability. The total porosity and microporosity of the mortar indicated that the TGP concrete had a denser microstructure than the other concretes. Air voids are also crucial for the strength and frost resistance of concrete [115], [116]. Multiscale pore structures, including air voids, capillary pores, and gel pore structures, suggest that the concrete maintained a good macro-performance. The TGP was inspected by the state in 2020, which revealed that the concrete was in normal service until now.
The phase assemblages, which are important for the microstructure, were not intensively studied in this study because of the complexity of dam concrete; however, a preliminary understanding of these can be obtained here. The granite aggregates occupied most of the space in the microstructure. Thus, the ITZs had a greater impact on the performance of the concrete. The structures, which existed widely at the boundary between the aggregate and paste, were as dense as the matrix. The partially reacted FA was evenly embedded in the hardened paste. However, FA particles were randomly clustered at the edges of the aggregates or in the matrix, resulting in porous regions. A similar porous microstructure owing to heterogeneous distribution was also reported in a previous study [99]. After 17 years, the MHC clinkers were almost completely hydrated to form C-S-H gels. Furthermore, the long-term reaction of FA consumed portlandite to form additional C-S-H gels that filled the microstructure. Although the reaction of siliceous FA proceeds relatively slowly [76], the reaction of the glass phase produces abundant C-S-H gels [94]. Therefore, more homogeneous gels with lower microporosity occupied most of the paste volume, thus improving the gel–space ratio described in Ref. [71]. Smaller FA particles were almost imperceptible in the BSE images, whereas larger particles remained almost intact, which revealed the heterogeneous reaction of FA [76]. Many residual FA particles in the microstructure [96] indicated the low reaction degree of FA, which is consistent with the experimental results and simulations from other studies [77], [96], [97], [117], [118], [119]. Relatively few crystalline phases were observed in the microstructures of the pastes. The crystalline phases in the microstructure mainly included Ms (or other AFm phases) [120], unexhausted portlandite (in concrete blended with 20% FA) [77] and partially hydrated periclases [23]. That is, the phase assemblage and homogenous distribution of the microstructure also reflected the high compactness and directly and indirectly improved the properties of the long-term dam concrete.
5.2. Impact of FA on long-term concrete
FA plays an important role in the long-term evolution of the microstructure [4]. Fig. 28 indirectly shows the contribution of FA to the long-term development of the microstructure. Increasing the FA content is more conducive to a continuous improvement in the performance of concrete. Concrete blended with the same FA content exhibited a higher final strength growth when the w/b ratio was higher. This indicates that the FA reaction and clinker hydration were no longer limited by the lack of space when the w/b ratio increased [73]. Both the phase assemblage and pore structure of the composites blended with FA changed significantly because of the reaction of FA [71], [94]. The pozzolanic reaction of FA consumes portlandite to form C-S-H gels. Further densification of the microstructure was promoted by the continuous formation of pozzolanic reaction products in all available pore spaces [76], [77]. The FA reaction products fill the capillary pores but barely grow in gel pores with a size of a few nanometers [73]. Hence, the gel pore distributions reflect the intrinsic information of C-S-H. The gel pore volume increased with an increasing FA content, which probably indicates the formation of additional gels because of the FA reaction. Moreover, FA significantly affects the capillary pore volume of concrete, especially in the later periods [91]. Fig. 29 shows the impact of the w/b ratio and FA content on the microporosity (10 nm–10 μm) of concrete. Without considering the FA content, the microporosity decreased with an increasing w/b ratio. When the substitution level of FA was constant, the variation in microporosity caused by the change in w/b ratio was not obvious. Fig. 29(b) shows the effect of the FA content on microporosity. The microporosity significantly decreased with an increasing FA content, regardless of the w/b ratio. This indicates that the effect of the w/b ratio on the micropore structure was distinctly weakened in the concrete during the very late period, but the contribution of FA to the microstructure was prominent. However, the w/b ratio significantly affected the total porosity, air void content, and macroscopic properties (Section 4.2). Excessive FA formed many loose regions (especially in Fig. 25(b)), resulting in a decrease in the compactness of the microstructure [99]. Blending more FA also reduces the reaction degree [94], [97]. Therefore, the influence of FA on the microporosity diminished when the dosage exceeded 30% (Fig. 29(b)). The RCPT results also revealed the importance of a 30% substitution level for FACB (Fig. 16).
FA consists of significant quantities of SiO2 and Al2O3 but little CaO; therefore, the blending of clinker with FA results in high amounts of aluminum-rich phases and C-S-H gels with a decreased calcium/silicon ratio [76], [77]. The uptake of aluminum by C-S-H led to the formation of C-(A)-S-H with a higher aluminum/silicon ratio. Although the chemical composition of the C-(A)-S-H gels of composites is very important for their structure and intrinsic properties [101], it was not analyzed further in this study. Additionally, an increase in Al2O3/SO3 results in an increase in AFm and a decrease in ettringite [120]. Hence, Ms was easily detected in the mineral compositions and microstructural images (Section 4.3.3).
The FA reaction degree can be calculated using various methods reported in previous study [119]; however, these quantitative techniques are suitable for pastes but not for concrete. The FA reaction degree in concrete was subjectively evaluated using SEM observations. The morphologies of hydrated FA are significantly different from each other because of variations in the intrinsic physical and chemical characteristics [102]. As shown in Fig. 30, the FA particles in the 17-year-old concrete were embedded in a C-S-H matrix covered by more flocculent hydrates. It is considered that the FA reaction degree is higher than that of four-year-old FA paste [92]. However, it was difficult to quantify the contribution of hydrated FA to the microstructure.
5.3. Impact of actual service environment
The effect of the curing mode on the performance of the concrete was investigated in this study. Concrete was exposed to the atmosphere for more than ten years to simulate the atmospheric environment of the dam concrete. The results of the mechanical properties, permeability resistance, and frost resistance demonstrated that natural curing in the atmosphere significantly affected the development of performance of the concrete, even without an aggressive aqueous environment [111]. Previous studies also revealed the impact of different complicated environments on the microstructure of concrete (Fig. 27). In outdoor environments, carbonation that results in carbonate precipitation is considered to have a major influence on the microstructure of concrete [69]. Additionally, the leaching caused by alternating wetting and drying cycles is a critical factor. Carbonation is typically thought to reduce the porosity of concrete [33]; therefore, an increase in the compressive strength was observed in the naturally-cured concrete, as shown in Fig. 8. However, in addition to an increase in strength, carbonation, and pore structure, natural carbonation also increased the number of coarse capillary pores [32], [35], [82], resulting in an increase in the diffusivity of concrete [34]. This enhances the negative effects of cyclic leaching on concrete [35], [106], which causes a decrease in the compressive strength and water permeability resistance (Fig. 8, Fig. 13). Furthermore, carbonation and leaching reduce the alkalinity of concrete [82], reducing the long-term reactivity of FA. The weakened contribution of FA to the strength increase during natural curing is shown in Fig. 28. The low porosity of each concrete sample demonstrates that the degradation of the microstructure by natural curing was not significant. Magnesium-rich gels were observed in the microstructure of the dam cores. A magnesium-rich gel is regarded as a mixture of C-S-H and M-S-H (magnesium silicate hydrate) owing to the different structures of these two silicates [121]. Carbonation and leaching decrease the pH of concrete, resulting in MgO easily forming M-S-H [105], similar to the blending of large amounts of FA [121], [122]. The change in the occurrence of magnesium in the microstructure, which is affected by the service environment of the dam, warrants further study.
Concretes in actual projects, such as tunnels, oil wells, and dams [41], [42], [43], [55], [57], continuously suffer from unexpected severe physical and/or chemical attacks. In the case of the Three Gorges Dam, the porosities of the concrete cores from the dam were higher than those of naturally-cured concrete, revealing that the concrete in the actual structure was subjected to more severe attacks. Both the porosity and phase assemblage changed, including the increase in carbonates, further decline of portlandite, and potential change of C-S-H compositions, which was also demonstrated by the mineralogical analysis and grayscale variations in the BSE images in Fig. 31. Additionally, the degree of degradation of dam concrete varies in various parts of the engineering field [41], [43], [56], and a significant reduction in microstructural compactness only occurred on the surface of the dam concrete from the exterior of the Three Gorges Dam. Continuous leaching caused by river water with traces of carbonates led to the dissolution of partial concrete hydrates on the exterior of the dam. The literature (Fig. 32 [106]) shows a similar degraded region owing to leaching, as shown in Fig. 26(c). The porous microstructure of the concrete in this region caused a significant increase in the porosity of the concrete on the external surface of the dam (Fig. 21). Although there was a certain degree of concrete degradation on the external surface of the dam due to carbonation and/or leaching, the performance loss of the entire concrete was negligible in terms of the spatial volume of the construction. However, the microstructure of concrete in particular parts of a dam is of particular concern, such as the engineering parts that are chronically or periodically exposed to river water [4], [41].
6. Conclusions
The long-term mechanical properties and durability of hydraulic concrete used in the TGP and the effects of natural and standard curing on the long-term performance of concrete were investigated in this study. Furthermore, both naturally-cured 17-year-old laboratory concrete and equivalent concrete cores drilled from the dam were used to investigate the microstructure of the concrete after a long period of hydration. Based on the results of this study, the following conclusions were drawn:
(1)Naturally-cured dam concrete exhibited a relatively good service performance. In terms of the evolution of mechanical properties over ten years, the long-term properties of the dam concrete for the TGP significantly exceeded the expected design values. The mechanical properties and durability of both the naturally- and standard-cured concretes continued to improve to different degrees after ten years. In particular, the compressive strength of dam concrete containing high amounts of FA (40%) exceeded 50 MPa with a relatively high w/b ratio of 0.55. There was no significant loss in durability for the dam concrete exposed to the natural environment for over ten years, except for the frost resistance. The number of freezing–thawing cycles for the naturally-cured dam concrete were significantly lower than those of the standard-cured concrete.
(2) The 17-year-old naturally-cured laboratory dam concrete was further analyzed. The RCPT results revealed the good service performance of the concrete, which still satisfied engineering requirements even after 17 years of natural exposure. The total porosity, air void structure, and micropore structure revealed that all the dam concretes possessed dense and homogenous microstructures, which is the origin of the excellent properties of the long-term dam concrete. Abundant and homogeneous C-(A)-S-H gels, crystalline phases (mainly AFm), and FA residues occupied the entire dense microstructural space in the pastes. Additionally, there was almost no porous ITZ between the aggregates and pastes.
(3) The microstructural analysis of the dam concrete revealed that most of the MHC clinker and FA with a fine particle size were reacted. Surprisingly, FA played a more important role in the microstructural evolution at later ages than the water–cement (w/c) ratio. Additionally, the performance enhancement attributed to FA appears to depend strongly on its replacement level. The microstructure and durability of concrete with > 30% FA can be remarkably improved. Furthermore, even after such a long period of hydration, periclase was still found in all systems, indicating that the periclase introduced during clinker calcination was not fully hydrated to form brucite.
(4) Despite the densified microstructure of the dam concrete and its excellent performance after 17 years of natural exposure, concrete drilled from the dam still exhibited a microstructure with decreased compactness, especially on the exterior of the dam. Except for the pore structures, the phase assemblage of concrete was also affected, including an increase in carbonates, a further decline in portlandite, and a potential change in the C-S-H composition. This can be mainly attributed to the change in hydrates due to carbonation and/or leaching caused by wind and water attacks. This phenomenon indicates that prolonged exposure to aqueous environments might cause the degradation of dam concrete, highlighting the need for further detailed research on the microstructure and performance evolution of concrete both in the laboratory and in the dam structure of the Three Gorges Dam.
Acknowledgments
The authors gratefully acknowledge the financial supports provided by the National Natural Science Foundation of China (U2040222, 52293431, and 52278259). The authors also wish to thank the test center of China Three Gorges Corporation for providing the concrete specimens and some data of environment, materials, and durability in this study.
Compliance with ethics guidelines
Chen Lyu, Cheng Yu, Chao Lu, Li Pan, Wenwei Li, and Jiaping Liu declare that they have no conflicts of interest or financial conflicts to disclose.
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