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Article

Study on the Effect of Water–Binder Ratio on the Carbonation Resistance of Raw Sea Sand Alkali-Activated Slag Concrete and the Distribution of Chloride Ions after Carbonation

by
Yan Wu
1,
Sixiang Kang
2,
Feng Zhang
3,
Haisheng Huang
4,
Haojie Liu
2,
Jianbin Zhang
1,
Hongze Li
5,
Weihong Li
6,
Zhou Zheng
1 and
Wenda Wu
2,*
1
Xiamen Municipal Engineering Design Institute Co., Ltd., Xiamen 361000, China
2
College of Advanced Manufacturing, Fuzhou University, Quanzhou 362200, China
3
CSCEC Strait Construction and Development Co., Ltd., Fuzhou 350015, China
4
College of Civil Engineering, Putian University, Putian 351100, China
5
Xiamen Municipal City Development and Construction Co., Ltd., Xiamen 361000, China
6
Dalian Municipal Design and Institute Co., Ltd., Dalian 116000, China
*
Author to whom correspondence should be addressed.
Buildings 2024, 14(7), 2027; https://doi.org/10.3390/buildings14072027 (registering DOI)
Submission received: 29 May 2024 / Revised: 25 June 2024 / Accepted: 28 June 2024 / Published: 3 July 2024
(This article belongs to the Collection Advances in Sustainable Building Materials and Construction)
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Abstract

:
The excessive extraction of river sand has led to significant ecological issues. Moreover, the environmental impact and resource demand of cement production have increasingly turned the spotlight on sea sand as a viable alternative due to its abundance and ease of extraction. Concurrently, alkali-activated binders, a novel type of low-carbon cementitious material, have gained attention for their low energy consumption, high durability, and effective chloride ion fixation capabilities. However, they are susceptible to carbonation. Introducing a controlled sea sand amount can raise the materials’ carbonation resistance, although carbonation may raise the concentration of free Cl− within the structure to levels that could risk the integrity of steel reinforcements by accelerating corrosion. In this context, the current study investigates sea sand alkali-activated slag (SSAS) concrete prepared with varying water–binder (W/B) ratios to evaluate its impact on flowability, mechanical strength, performances, and chloride ion distribution post-carbonation. The results demonstrate that the mechanical property of SSAS concrete diminishes as the water-to-binder ratio increases, with a more pronounced reduction observed. The depth of carbonation in mortar specimens also rises with the W/B ratio, whereas the compressive strength post-carbonation initially decreases before showing an increase as carbonation progresses. Furthermore, carbonation redistributes chloride ions in SSAS, leading to a peak Cl− concentration near the carbonation front. However, this peak amplitude does not show a clear correlation with changes in the W/B ratio. This study provides a theoretical foundation for employing sea sand and alkali-activated concrete.

1. Introduction

The accelerated expansion in construction industries worldwide has catalyzed economic growth but has simultaneously contributed to various environmental issues. Cement production, in particular, is a significant environmental concern because of its extensive consumption of natural resources and the substantial emissions of carbon dioxide it generates during production. It is calculated that cement production accounts for approximately 8% of global carbon emissions annually, presenting a critical challenge to the sustainable development of China’s construction sector [1]. In response, geopolymers have emerged as a sustainable alternative, attracting broad interest. These materials utilize industrial waste, thus reducing reliance on traditional cement and decreasing the carbon footprint while recycling industrial by-products [2,3,4,5]. Geopolymers are noted for their strong adhesive properties, rapid strength development, and superior durability, confirming their potential as an effective substitute for conventional cement through extensive research [6].
Meanwhile, as the global demand for sand in concrete production continues to climb, the gap between the supply and demand for river sand is widening. The overexploitation of river sand not only leads to ecological degradation and resource scarcity but also consumes significant energy—approximately 0.083 gigajoules per ton [7]. The extraction processes have substantial negative impacts on river ecosystems, affecting navigation and flood prevention efforts, thereby accelerating the transition towards sea sand as an alternative [8]. Previous research indicates that concrete made with sea sand can display superior mechanical properties [9,10,11]. This improvement is attributed to the similar source and mineral composition of sea sand (SS) compared to river sand (RS). Additionally, SS typically features low mud content and excellent particle size distribution, making it suitable for concrete production after undergoing desalination and demineralization treatments [12,13]. Newman [14] concluded that the key difference between sea sand and river sand lies in the presence of broken shell materials and intact gastropod particles. Due to the basic chemical composition of shell particles (CaCO3), sea sand has a higher density. The shell fragments are sturdy and durable, which can reduce porosity. Most researchers believe that the early and long-term compressive strengths of sea sand concrete are slightly higher than those of ordinary concrete [15,16]. Additionally, Liu found that sea sand concrete exhibited better flexural strength than ordinary concrete [14]. Various studies have also investigated the steel reinforcement corrosion of sea sand concrete [17], as well as the impact of the shell content on its mechanical properties [18], confirming its excellent performance. These findings collectively demonstrate the feasibility of using sea sand as a substitute for river sand.
While employing SS as a fine aggregate presents a viable alternative to alleviate the pressure on river sand extraction, its inherent high salt content poses challenges for its widespread application. The presence of salts, particularly chlorides in untreated SS, can compromise the durability of building structures by promoting reinforcement corrosion and accelerating carbonation, especially in marine and coastal settings [19]. Desalination has emerged as a promising solution to diminish salt content in SS, yet it is not without drawbacks. The process requires substantial water and energy resources, incurring high operational costs which pose economic challenges for its widespread adoption, especially in seaside cities and island areas where desalination facilities may be less accessible. Additionally, the conventional method of desalinating sea sand by soaking it in freshwater demands large volumes of water. Amid these challenges, there is increasing research interest in untreated sea sand and addressing the durability issues associated with its use in construction. For instance, Wei’s experiments [20], which incorporated fly ash and a blast furnace into sea sand concrete, had limited success in enhancing the passivation film’s density. Conversely, Jiang’s [21] study indicated that sea sand concrete with a 20% fly ash mix showed promising mechanical properties and durability, noting that chloride ions positively influenced the early strength of the concrete but adversely affected its later stages. Additionally, Zhu [22] explored the introduction of gaseous SiO2 into concrete to modify the formation mechanism of Friedel’s salt, aiming to reduce Cl− permeation. However, this approach proved economically inefficient and challenging to implement on a large scale in engineering projects, highlighting the complexities of developing viable solutions for the utilization of SS in construction.
It is worth nothing that compared to traditional concrete, alkali-activated concrete has a better ability to bind Cl− [23,24,25]. This attribute effectively mitigates the corrosion risks linked to the Cl− in sea sand, thus facilitating its broader adoption in engineering practices. Research focusing on alkali-activated systems in marine environments has revealed that these materials often exhibit an increase in strength even after prolonged exposure to seawater [26]. Consequently, they present a promising alternative for construction in harsh marine environments.
Indeed, the durability of concrete in marine environments is influenced not only by chloride ions but also by the synergistic effects of these ions and carbonation. Carbonation alters the microstructure of concrete, which can accelerate the Cl− penetration. This process also reduces the materials’ pH, compromising the stability of chloride ion binding. It results in the breakdown of calcium silicate hydrate and Friedel’s salt, resulting in the release of previously bound Cl− [27,28,29]. As a result, the concentration of Cl− in the pore solution increases [22,30], which in turn affects the distribution of these ions [31] and directly correlates with the steel corrosion in concrete [27,32]. For alkali-activated concrete, which typically exhibits lower alkalinity and a lower calcium hydroxide (Ca(OH)2) content, there is generally a deeper penetration of carbonation compared to traditional silicate cement [33]. However, some studies indicate that the introduction of Cl− from sea sand may alter the concrete porosity, influencing the transport behavior of CO2 [34] and potentially enhancing the concrete’s carbonation resistance [35,36]. Therefore, researching the complex interaction between chloride ions and carbonation in alkali-activated concrete is essential. This topic deserves thorough investigation to better harness the properties of these materials for improved performance in challenging environments like marine settings.
Our preliminary research has shown that the synergistic combination of alkali-activated concrete and sea sand can be highly effective [37]. Considering the shell content, particle size distribution, and the comprehensive impact of active ions in sea sand on the performance of the geopolymer, we determined the optimal sea sand content to be 30%. At this content, the geopolymer exhibits the lowest porosity, resulting in excellent mechanical properties and carbonation resistance. But one challenge faced in the application of alkali-activated concrete is its tendency to set rapidly, particularly at lower water–binder (W/B) ratios. This rapid setting can lead to the poor flowability of the mortar, complicating the molding process and potentially compromising the uniformity of the mortar mix, which in turn impacts the overall properties of the material. To address these issues, this study has prepared various formulations of sea sand alkali-activated slag (SSAS) concrete using slag, sea sand, and sodium silicate, varying the W/B ratio. The focus of this article is to explore how the W/B ratio affects the post-carbonation morphology, carbonation depth, compressive strength, and chloride ion distribution in SSAS. We aim to elucidate the mechanisms involved through detailed microanalytical methods such as XRD, TG-DTG, and MIP. This investigation fills a critical gap in understanding the dual impact of intrinsic Cl− erosion and carbonation on the SSAS. The insights gained will contribute to a deeper understanding of rebar corrosion mechanisms in alkali-activated sea sand concrete. Moreover, this research provides essential theoretical and technical insights for structural durability in marine engineering environments, aiming to enhance the longevity and structural integrity of such constructions.

2. Materials and Methods

2.1. Materials

2.1.1. Slag

The slag used in this study is S95 grade granulated blast furnace slag. The basic parameters are detailed in Table 1, while the results from the XRF chemical composition analysis are presented in Table 2. Additionally, the results of the XRD analysis of the slag are depicted in Figure 1. The results reveal that the predominant component of the slag is glassy, with quartz and calcite being the primary crystalline phases identified. This composition suggests specific behaviors in the slag’s reactivity and its potential effectiveness in alkali-activated formulations.

2.1.2. Sand

The sand utilized in our experiments includes natural sea sand from Zhangpu County, Zhangzhou City, Fujian Province, and natural river sand from Fujian Province. Notably, these sands were not dried prior to their basic characterizations. The fineness modulus of the SS is measured at 2.83, while the RS is measured at 3.03. The moisture content, mud content, chloride ion content, and shell content of these sands were evaluated according to the standards outlined in GB/T 14684-2011 [38], and all parameters met the required specifications, as detailed in Table 3. Before being used in the experiments, the sand was dried under sunlight to ensure consistency in the testing material.
The fineness and particle size distribution of sea sand and river sand were determined according to the JGJ52-2006 standard [39]. The cumulative sieve residue percentages of the sand are shown in Table 4. The nominal diameters of the sieve openings are 5.00 mm, 2.50 mm, 1.25 mm, 0.63 mm, 0.315 mm, and 0.16 mm, with the cumulative sieve residue percentages of the sand being A1, A2, A3, A4, A5, and A6, respectively. The fineness modulus of SS was calculated to be 2.83, classifying it as medium sand, while the fineness modulus of RS was calculated to be 3.03, also classifying it as medium sand.

2.1.3. Alkali Activator

The alkali activator in this study is anhydrous Na2SiO3, which is granular with a modulus of 1.0. The composition of the solid activator includes 50.82% Na2O and 49.18% SiO2. The preparation and selection of materials are fundamental for ensuring the reliability and reproducibility of the experimental results.

2.2. The Preparation and Mixing Ratio of Sample

The alkali-activated cement was prepared using a mechanochemical approach, involving the mixing and milling of slag with an alkali activator. This process utilized a WZM-15 × 2 ball mill machine (in Figure 2 and Figure 3), which features a 15 L capacity and double-barrel configuration, with an inner diameter of 26 cm for the ball milling jar. The ball mill uses steel balls of 10, 20, and 30 mm diameters, totaling 15 kg, in a mass ratio of 1:1:1.
The ball milling procedure began by mixing the dried slag with the alkali activator in a specific ratio and transferring the mixture into the ball mill jar loaded with the steel balls. The ratio of the total mass of the steel balls to of the mixture was maintained at 7.5:1, with the total mass of the alkali activator and slag being 2 kg. The sealed jar was then placed in the ball mill and subjected to high-speed milling at 85 revolutions per minute for one hour. Post-milling, the resultant powder was sifted through a screen to ensure a uniform particle size and stored in a sealed bag, ready to be used as the cement component for further experiments. Subsequently, the prepared cement was combined with sand in a mixer to produce the SSAS samples. Test group fit examples are shown in Table 5.
The milling parameters were determined based on preliminary experimental results. When the rotational speed is too low, the balls inside the mill are lifted to a small height, resulting in low impact force. Conversely, if the speed is too high, the balls enter a centrifugal state, causing the grinding to stop. If the milling time is too short, the particle fineness remains large and the activity is low; if the time is too long, particle agglomeration occurs, reducing activity [40]. Through initial testing, we optimized a grinding speed of 85 r/min and a duration of 1 h, which resulted in the geopolymer achieving optimal performance. The ball-to-material ratio was determined based on the required raw material quality and overall milling efficiency.

2.3. Testing and Analysis

2.3.1. Fluidity Test

The flowability of the mortar is tested in accordance with the standard GB/T 2419-2005 [41].

2.3.2. Mechanical Property Test

The methods for testing the mechanical properties of the mortar are carried out following the standard GB/T 17671-1999 [42]. The mortar is molded into a triple form with dimensions of 40 mm × 40 mm × 160 mm. After demolding, the specimens are transferred to a standard curing room maintained at 20 ± 2 °C with a relative humidity over 95%. They remain there until they reach predetermined ages of 3, 7, 14, and 28 days. After curing, their flexural and compressive strengths are evaluated using a DYE-300 type integrated tester (Produced by Haoqiang Technology, Nanjing, China). Values that exceed ±10% of the average are discarded and the measurements are performed again.

2.3.3. Carbonation Depth

The carbonation test for the mortar specimens with dimensions of 70.7 mm × 70.7 mm × 70.7 mm, is conducted following rigorous protocols to ensure accuracy and replicability. After undergoing standard curing for 26 days (at 20 ± 2 °C with relative humidity over 95%), the specimens are taken out and placed in an oven set to 40 °C, where they undergo a drying process for 48 h. Once dried, one face of each specimen is left exposed to carbonation, while the remaining five faces are sealed with paraffin wax to prevent unintended exposure to carbon dioxide. Following exposure, the specimens are transferred to a carbonation chamber and maintained under controlled conditions with 3 ± 0.5% CO2, 20 ± 2 °C, and a relative humidity of 70 ± 5%. This setup adheres to the guidelines specified in GB/T 50082-2009 [43].
At predetermined carbonation intervals of 3, 7, 14, and 28 days, the samples are carefully cut open using a precision cutting machine. After cutting, the newly exposed surfaces are dried and subsequently sprayed with a phenolphthalein solution to evaluate the carbonation depth. The solution used for this test consists of 1 g of phenolphthalein and 99 g of anhydrous ethanol, creating a 1% solution. When applied to the freshly cut face of the specimen, this solution reacts with the concrete. Areas of the concrete that have not undergone carbonation will turn purple upon contact with the phenolphthalein, while carbonated areas remain unchanged. The carbonation depth is indicated by the distance from the specimen’s outer surface to the point where the color change ceases. This measurement is taken using a scale ruler, providing a quantitative measure of the extent of carbonation at each testing age. This method allows for a clear and precise determination of how deeply CO2 has penetrated the mortar, which is crucial for understanding the material’s durability and resistance to environmental conditions.

2.3.4. Compressive Strength after Carbonization

The assessment of the compressive strength of mortar specimens post-carbonation at various stages—3, 7, 14 and 28 days—follows the ISO method. A YAW4206 type universal testing machine (Produced by Wance, Shenzhen, China) is used to measure compressive strength. This advanced equipment ensures precision by applying a uniform load at 5000 N/s until the specimen fails, at which point the compressive strength data are recorded.

2.3.5. Chloridion Concentration Test

For specimens that have undergone carbonation and curing for 28 days, a detailed analysis of the chloride ion concentration is conducted. This involves collecting powder from the specimens layer by layer. The grinding depth is precisely adjusted to reach a total of 25 mm, with samples taken every 2 mm for the first 10 mm of depth and every 5 mm thereafter, up to a 25 mm depth, resulting in eight distinct layers of powder. This process yields eight distinct powder layers. Each powder sample is then sifted through an 80 μm sieve and dried until it reaches a constant weight, making it ready for testing the Cl− concentration.
The acid-soluble method is used to determine the total Cl− concentration. This process, as detailed in the relevant literature, includes steps such as boiling, dissolving, and filtering to produce a test solution of chloride ion concentration. Each powder layer’s total chloride ions are extracted using HNO3, and the consumption of silver nitrate AgNO3 is measured by titration for each sample to calculate the chloride ion concentration [44].

2.3.6. X-ray Diffraction Test

Samples are collected using a concrete profile grinding machine and subsequently sieved through an 80 μm standard cement sampling sieve. These samples, extracted uniformly from a depth of 4–6 mm within the carbonated specimens, are placed in isopropanol to halt any further hydration. The isopropanol is replaced every 24 h, with a total of three replacements occurring. After seven days, the wet samples are transferred into labeled aluminum boxes and placed in a vacuum drying oven for drying. The analysis of hydration reactions and types of crystals formed post-carbonation is conducted using a Miniflex 600model X-ray diffractometer (Produced by Japan Mechanics Rigaku, Tokyo, Japan). This equipment is set to scan from 5° to 80°, adjusting the step size to 0.02° per second and a scanning speed of 10° per minute. This setup ensures a detailed examination of the crystalline structures that develop because of the carbonation process.

2.3.7. Thermogravimetric Analysis

The TG analysis sample preparation follows the same methodology as that used for the XRD analysis sampling. The analysis utilizes a TGA/DSC3+ thermogravimetric analyzer (Produced by Mettler Toledo, Greifensee, Switzerland). During the experiment, the samples are heated at a rate of 10 °C/min, under a gas flow of 20 mL/min. Standard cured samples are heated from room temperature up to 800 °C and held for 20 min. In contrast, carbonated cured samples are heated to 1000 °C. Nitrogen is employed as the protective gas throughout the process. This setup automatically records the TG/DSC curves of the samples.
Some substances within the carbonated mortar may undergo dehydration or decomposition reactions when heated. The TG data collected allows for the calculation of the first derivative, known as the DTG. By analyzing both the TG and DTG curves, qualitative and quantitative assessments of specific substances in the samples can be performed, providing insights into the chemical changes occurring during the heating process.

2.3.8. Mercury Intrusion Porosimetry

Concrete cutting machines are employed to precisely slice mortar specimens into sections approximately 5 mm thick. Following this, a small hammer is utilized to gently chip off pieces from these slices. For the purposes of MIP, fragments that are square or circular in shape and approximately 3–5 mm in diameter are carefully selected.
After the selected samples have undergone hydration cessation and drying, their pore structures are analyzed using an AutoPore 9500 pressure porosimeter (Produced by MicroActive, Norcross, GA, USA). The porosimeter operates within a testing pressure range from 0.10 to 61,000.00 psia, which allows for a detailed examination of the samples’ microstructural features. The mercury intrusion test conducted provides critical data on porosity and the distribution of pore sizes, offering insights into the material’s permeability and structural integrity.

3. Results and Discussion

3.1. Flowability

The flowability of the SSAS at different W/B ratios is shown in Figure 4. It was found that at a water–binder ratio of 0.39, the flowability is only about 163 mm, which can further decrease during casting due to rapid setting, adversely affecting the molding of specimens. This also impacts the uniformity of the specimens, resulting in a surface and internal structure that is porous and unevenly distributed. The flowability of the mortar increases with the W/B ratio; so, subsequent experiments were conducted, using a 0.43 W/B ratio, 0.47 W/B ratio, and 0.51 W/B ratio as variables. At a W/B ratio of 0.43, the flowability reached about 210 mm, and the mortar exhibited good workability throughout the casting process, resulting in smoothly molded specimens. When the W/B ratio was further increased to 0.51, although the flowability was higher, exceeding 270 mm, there were no issues with segregation or bleeding, and the specimens were normally formed.

3.2. Mechanical Property

Figure 5 shows the effects of different water-to-binder ratios under standard curing conditions on the mechanical property of SSAS concrete. It can be observed from the figures that under standard curing conditions, both the flexural and compressive strengths of mortar specimens decrease with increasing water-to-binder ratios at different ages. After 28 days of standard curing, compared to the W0.43 group, the flexural strength of the W0.47 group decreased by 8.33% and the compressive strength decreased by 5.37%; the W0.51 group’s flexural strength decreased by 16.67% and the compressive strength decreased by 14.56%. The reduction in compressive strength after 28 days of standard curing increases with the water-to-binder ratio. As the W/B ratio increases, the flexural and compressive strengths of the mortar gradually decrease. This reduction can be attributed to the following factors: a higher water content can dilute the alkaline activator, reducing its effectiveness and slowing the hydration process. This decrease in the hydration products formed increases porosity and ultimately lowers the strength of the mortar [45]. Additionally, an increased water–binder ratio after the polymerization reaction leads to higher residual moisture within the material. This excess water remains trapped, forming water pockets or bubbles, which negatively affect the material’s strength performance [46].
This is because an increase in the water–binder ratio lowers the initial concentration of silicate ions and alkalinity, reducing the quantity of hydration products formed, leading to an increase in porosity and a consequent decrease in the strength of the mortar [22].
The flexural strength of groups W0.43 and W0.47 essentially completes its development by 14 days, while group W0.51 continues to grow slowly after 14 days; at the same time, compared to group W0.43’s compressive strength, the compressive strength of groups W0.47 and W0.51 gradually narrows the gap with W0.43 as the curing age increases. Moreover, when the water-to-binder ratio is lower, the strength development of the mortar specimens is concentrated in the early stages of curing. This is because the increased amount of water may have a diluting effect, leading to a delayed hydration process and a relatively slow strength development [47].

3.3. Properties after Carbonization

3.3.1. Carbonation Depth

Figure 6 shows the color rendering of alkali slag mortar specimens with different water–binder ratios at 3 d, 7 d, 14 d, and 28 d of carbonation. The carbonation depth of the samples can be measured according to the color they present. As the carbonation days increase, the purple areas on the surface of the specimens gradually decrease, indicating an increasing degree of carbonation. The specimens carbonated for 28 days have the smallest purple areas, indicating the highest degree of carbonation. At the same carbonation age, specimens with a higher water–binder ratio (W0.51) show a faster carbonation rate. The higher water–binder ratio leads to increased porosity in the concrete, allowing carbon dioxide to penetrate and react more easily. Consequently, the degree of carbonation is higher, and the reduction of the purple areas is more pronounced.
Figure 7 displays the coloration images of SSAS specimens at various W/B ratios at 3, 7, 14, and 28 days of carbonation. The depth of carbonation can be determined based on the color presented by the samples. The figure clearly illustrates that as the W/B ratio rises, the depth of carbonation in mortar specimens also increases across various carbonation ages, with the rate of increase becoming more pronounced at higher ratios. This trend indicates a corresponding decrease in the carbonation resistance of the SSAS. Specifically, after 28 days of carbonation, the carbonation depth in the W0.47 group increased by 7.31% compared to the W0.43 group, and in the W0.51 group, it increased by 20.73%.
This escalation in carbonation depth continues as the carbonation process progresses, although the rate of change tends to flatten over time. Geopolymers, which are typically activated by alkaline solutions, are affected by increases in the W/B ratio because the dilution of these activators reduces their effectiveness. This diminished effectiveness leads to less efficient polymerization processes and, consequently, a reduced generating of C-(A)-S-H gel. This gel is crucial for resisting environmental erosion, including carbonation. Moreover, an increase in the W/B ratio also elevates the amount of free water in the paste. Once this water evaporates, it generally results in increased porosity within the material. Higher porosity allows for easier penetration of carbon dioxide and other gases, further exacerbating the carbonation. This phenomenon underlines the critical balance needed in the W/B ratio to optimize the durability and environmental resistance of geopolymers in construction applications [48]. Similar to SSAS concrete, increasing the water–binder ratio also enhances the carbonation reaction and CO2 absorption in Portland cement. However, traditional Portland cement concrete typically has a higher Ca(OH)2 content, which acts as a buffer for carbonation by consuming carbonate ions [49].
As the carbonation process progresses, the carbonation depth increases with extended exposure time. During the initial 7 days, the carbonation depth of all specimens increased rapidly. This rapid increase can be attributed to the direct contact of SSAS surfaces with CO2, allowing CO2 to penetrate and quickly react with the gel inside the concrete. From 7 to 14 days, the carbonation depth continued to increase, but at a slower rate. This indicates that the formation of the surface carbonation layer began to hinder further CO2 penetration [50]. In the last 14 days, the rate of increase in carbonation depth slowed down even further. This could be due to the surface carbonation layer becoming denser, providing a stronger barrier to CO2 penetration.

3.3.2. Compressive Strength after Carbonization

The compressive strength of SSAS concrete after carbonation is shown in Figure 8 at varying W/B ratio and carbonation ages. After 3 days of carbonation, the compressive strength of the sample exhibited a decline compared to the pre-carbonation sample, influenced by different W/B ratios. This decline is attributed to several factors. First, size discrepancies between standard cured specimens and carbonation test specimens create a size effect [51]. Additionally, carbonation lowers the Ca2+ concentration in pore solution, leading to the decalcification of the hydration product C-S-H [52]. This reaction not only decreases the pH, potentially weakening the structural integrity of the C-S-H gel [53]—the primary strength contributor of cement—but also transforms it into less robust gels like silica gel and alumina gel, with a lower Ca/Si ratio. Consequently, this softens the paste, diminishes the bonding strength, and reduces the compressive strength [54]. Early carbonation stages also release bound-free water from CH, causing drying and surface shrinkage cracks, which are more pronounced in samples with higher water–binder ratios [55,56]. Specifically, within the initial 0–3 days, compressive strength decreased by 6.8%, 9.5%, and 12.8% in the W0.43, W0.47, and W0.51 groups, respectively.
Beyond 3 days, extending the carbonation age generally enhances the compressive strength, although the rate of increase from 3 to 28 days diminishes with higher water-to-binder ratios. For instance, after 28 days of carbonation, the compressive strengths in the W0.43, W0.47, and W0.51 groups rose by 12.83%, 9.92%, and 7.21%, respectively. This improvement results from additional carbonation forming CaCO3, which fills the pores and reduces the total porosity, thereby densifying the microstructure [57,58]. However, as the W/B ratio ascends, the compressive strength of mortar after carbonation progressively declines at different ages. The extent of this decrease amplifies with a higher W/B ratio. After 28 days, the compressive strengths in W0.47 and W0.51 groups dropped by 10.54% and 24.05%, respectively, compared to the W0.43 group, a reflection of a lower degree of hydration and higher porosity [45].

3.3.3. The Distribution of Chloridion under Carbonation

Figure 9 depicts the chloride ion concentration profile in SSAS concrete across various water-to-binder ratios. The figure shows that Cl− concentration is initially lowest at the specimen surface, increases to a peak, and then diminishes deeper within the specimen. Despite the expectation of a uniform chloride ion distribution due to the thorough mixing of sea sand during the sample preparation, carbonation significantly alters this distribution. After carbonation, chloride ions congregate around the carbonation front—marked by a phenolphthalein color change—at the boundary between carbonated areas and uncarbonated areas [59,60]. This clustering arises because carbonation liberates some bound Cl−, generating the free Cl− in the carbonated zone and establishing a concentration gradient [37]. This gradient drives the diffusion of Cl− from the carbonation front towards the non-carbonated interior, leading to a localized Cl− peak at this interface. The peak concentration of chloride ions near the carbonation front indicates a higher risk of steel reinforcement corrosion after carbonation. This finding suggests that although SSAS concrete can initially bind chloride ions effectively, carbonation can release these ions, creating localized concentrations that accelerate corrosion. Therefore, for SSAS concrete structures exposed to marine environments, additional protective measures may be necessary to ensure long-term durability [61].
In groups W0.43, W0.47, and W0.51, the initial chloride ion concentrations are 0.0285%, 0.0282%, and 0.0279%, respectively. Post-carbonation, the peak concentrations observed are 0.0309%, 0.0290%, and 0.0280% for these groups, indicating increases from their initial levels. This phenomenon highlights the potential for enhanced Cl− induced corrosion at the carbonation front compared to initial conditions. Despite similar initial concentrations, a rise in the W/B ratio leads to lower initial surface chloride concentrations but higher peaks. This pattern suggests that with higher water–binder ratios, which typically exhibit incomplete hydration and increased matrix porosity, more bound chloride ions are released and migrate inward. This porosity not only allows for more chloride ion release but also accelerates their transport, consequently shifting the peak concentration inward and modifying its intensity. In the higher W/B ratio groups, the Cl− concentration in the non-carbonated area notably exceeds that in groups with lower ratios, indicating more extensive chloride ion migration. This dynamic suggests that in samples with a higher W/B ratio, the risk of rebar corrosion might be more pronounced due to elevated internal chloride concentrations.

3.4. XRD

Figure 10 presents the XRD spectra of SSAS specimens with a varying W/B ratio after 28 days of standard curing. The spectra reveal that although the hydration product types remain consistent across different ratios, the intensity of their peaks varies. The primary hydration product, C-(A)-S-H gel, is evident through a broad, diffuse peak between 15° and 40° [62]. Notably, as the W/B ratio increases, the diffraction peak height of C-(A)-S-H gel diminishes, suggesting a declination in its formation. This reduction correlates with a gradual deterioration in the mechanical property of mortar, as the hydration products that typically enhance the internal pore structure and provide strength are less prevalent in higher water-to-binder ratio samples. Additionally, Gismondine (CaAl2Si4O12·H2O) forms a diffraction peak around 21°, which also shows a decline in intensity with increasing water-to-binder ratios. The presence of quartz peaks originates from the sea or river sand used in the mortar, while calcite and aragonite peaks, stemming from seashell components in the sand, also feature in the spectra.
Figure 11 displays the XRD spectra of SSAS mortar after 28 days of carbonation at different water–binder ratios. Similar to the non-carbonated specimens, no new phases emerge with increasing water-to-binder ratios in the carbonated samples. However, the intensity of the calcite peaks grows. This increase indicates that specimens with higher alkali equivalents, which exhibit a denser pore structure, hinder CO2 penetration, affecting the rate and extent of carbonation. The peak near 28° sharpens over time, reflecting the ongoing carbonation of C-(A)-S-H gel and the generation of CaCO3. Despite a denser pore structure limiting CO2 entry, which theoretically should reduce CaCO3 formation, the characteristic peak of CaCO3 becomes more pronounced. Scholars note that a CO2 concentration around 3% significantly accelerates carbonation, primarily yielding calcite [63]. Nonetheless, the XRD spectra also identify vaterite and aragonite, with an endothermic peak observed at approximately 450 °C in the thermal analysis. These findings suggest that during early carbonation stages, some CaCO3 produced from the decalcification of C-(A)-S-H initially exists in a less stable form, poorly crystallized vaterite and aragonite [64,65], which eventually convert to calcite as carbonation advances [66].
Figure 12 illustrates the SSAS XRD at different W/B ratios across various carbonation ages. The spectra highlight that CaCO3 is present, primarily in the form of calcite at diffraction angles of 28°. This peak is not evident in the samples before carbonation but begins to appear after 3 days of carbonation curing. Because this peak overlaps with the broad peak of C-A-S-H gel, it is difficult to determine the exact height of the peak. The specific amount of CaCO3 formed will be further analyzed in the TG analysis below. The formation of calcite not only fills the internal pores of the mortar but also significantly contributes to its structural strength, thus enhancing the compressive strength of samples as carbonation progresses. This beneficial relationship between carbonation age and mechanical property is primarily due to the effective pore-filling and strengthening properties of calcite.
In addition to calcite, the carbonation products include smaller quantities of aragonite, visible at 27° and 46°, and vaterite at 50°. Research indicates that the CaCO3 formed from the carbonation of Ca(OH)2 crystallizes into relatively complete calcite. During carbonation, part of the CaCO3 formed from the decalcification and carbonation of C-S-H gel in the mortar specimens appears as aragonite and vaterite, which are less stable and have poorer crystallinity compared to calcite. The decomposition temperature of these forms is lower than that of calcite [64,65]. The TG analysis revealed endothermic peaks of aragonite and vaterite around 450 °C. Compared to calcite, the formation of unstable aragonite and vaterite in the early stages of carbonation has a weaker effect on strength improvement. The XRD spectra indicate the quantity of aragonite rises with the advancement of the carbonation age. However, the changes in vaterite are indiscernible due to its peaks overlapping with those of calcite and quartz. The C-S-H gel peak in the mortar specimens weakens as carbonation progresses, and the diffraction peaks gradually become sharper. This is because the carbonation process causes CO2 to react with some of the hydration products to form CaCO3.

3.5. TG

The TG-DTG analysis is highly sensitive to the mass changes associated with the decomposition of specific phases and can provide valuable quantitative data on hydration and carbonation products. However, this technique has limitations, including the potential overlap of thermal events and the impact of sample preparation on accuracy. Despite these limitations, TG-DTG remains a useful tool for characterizing phase composition and understanding the thermal behavior of SSAS concrete.
The TG curve, delineated by absorption peak positions and subsequent material decomposition, can be divided into five distinct temperature ranges: below 150 °C, 150–400 °C, 400–500 °C, 500–800 °C, and above 800 °C. The 25–150 °C range corresponds to the first segment of mass loss, associated with the evaporation of free water and some partial dehydration of the C-S-H gel layers [67,68]. The 150–400 °C range corresponds to the second segment of mass loss, characterized by the loss of bound water and hydroxyl groups from hydration products such as C-S-H and C-(A)-S-H [40,69]. The interval between 400 and 500 °C represents the third phase, involving the dehydration of silica gel. In this stage, unstable aragonite decomposes into calcite, but this does not affect mass loss. The 500–800 °C range signifies fourth segment of mass loss due to the CaCO3 decomposition from carbonation.
Figure 13 and Figure 14 display the TG-DTG curves and mass loss graphs for the SSAS after 28 days of standard curing. Upon heating to 800 °C, the total mass losses for the W0.43, W0.47, and W0.51 groups are 7.88%, 6.91%, and 6.75%, respectively. The first decline peak appears around 87 °C, representing the evaporation of free water in the specimens, with the mass loss between 25 and 150 °C for the W0.43, W0.47, and W0.51 groups being 2.55%, 2.27%, and 2.20%, respectively. A second decline peak appears around 365 °C, where the mass loss from 150 to 400 °C is attributed to the loss of bound water in hydration products, with respective losses of 2.26%, 2.01%, and 1.94% for the groups. This suggests that the amount of hydration products in the mortar specimens decreases with increasing W/B ratios, leading to increased porosity and reduced compressive strength under standard curing, and relatively weaker carbonation resistance. A third decline peak appears around 491 °C, corresponding to the dehydration of silica gel formed from the NaSiO3 reaction with SiO2 in the slag. With an increasing W/B ratio, the amount of silica gel in the mortar specimens decreases. At around 636 °C, a fourth decline peak appears, corresponding to the decomposition of CaCO3. Given that the measured carbonation depth of standard-cured 28-day specimens is zero, indicating that no carbonation reaction occurred, the observed weight loss is attributable to the shells’ decomposition from sea sand.
Figure 15 and Figure 16 show the TG-DTG curves and mass loss graphs for the SSAS after 28 days of carbonation curing. A second decline peak appears around 275 °C, earlier than the temperature under standard curing conditions, with the mass losses between 150 and 400 °C for the different groups after carbonation being 2.36%, 2.06%, and 2.08%, respectively. This is because the quantity of hydration products is highest, impeding CO2 transport and enhancing carbonation resistance. Compared to samples before carbonation curing, the reduction in hydration products after carbonation curing is minimal, as hydration reactions continue with carbonation. A third decline peak appears around 451 °C, where aragonite decomposes into calcite due to heating, but the decomposition of aragonite does not cause mass loss [64], and aragonite’s presence was also found in the XRD diffraction peaks; thus, the temperature range of 400–500 °C corresponds to the dehydration of silica gel, with mass losses of 2.12%, 1.70%, and 1.49%, respectively, decreasing with an increasing water–binder ratio. The main reason is the decalcification reaction of the C-S-H gel under carbonation, which increases the silicon content, and the amount of silica gel after carbonation decreases with an increasing water-to-binder ratio. When the temperature reaches about 650 °C, a fourth decline peak appears; the temperature range of 500–800 °C corresponds to the decomposition of CaCO3, with mass losses of 2.92%, 3.00%, and 3.57%, respectively. Under carbonation curing conditions, the CaCO3 content shows a significant increase compared to standard curing, indicating that a substantial amount of CaCO3 forms under carbonation conditions, gradually filling the pores and reducing the porosity; thus, the compressive strength after carbonation shows an increasing trend. And at lower water–binder ratios, this part of the mass loss is lower, indicating a stronger carbonation resistance.
Figure 17 shows the mass loss rate of the SSAS mortar with a W/B ratio of 0.43 at different carbonation ages. After carbonation, the mass losses of each group of specimens between 150 and 400 °C were 2.45%, 1.91%, and 2.36%, respectively, showing a trend of first decreasing and then increasing with age. This indicates that the amount of C-(A)-S-H gel initially decreases and then increases. The mass losses between 400 and 500 °C were 0.87%, 1.58%, and 2.12%, respectively, indicating a significant increase in silica gel content. In the early stages of the carbonation reaction, C-S-H gel undergoes decalcification, forming silica gel, which cannot provide a sufficient strength support for the mortar. This causes the paste to soften, the bond strength to decrease, and the overall strength to decline [54]. The temperature range of 500–800 °C corresponds to the decomposition of CaCO3, with mass losses of 1.49%, 2.68%, and 2.92%, respectively. Compared to standard curing conditions, the CaCO3 content under carbonation curing conditions significantly increases, indicating that a large amount of CaCO3 is formed, gradually filling the pores and thus reducing porosity. This leads to an increasing trend in compressive strength during the mid to late stages of carbonation.

3.6. Pore Structure

The mechanical property and durability of the SSAS is significantly influenced by the nature of the specimen’s pore structure, and variations in these characteristics can more clearly explain the trends in macroscopic performance [70,71,72]. The academician Wu Zhongwei has categorized these pores based on their size and impact on the concrete [73]. Harmless pores that are smaller than 20 nm in diameter do not affect the concrete’s strength or durability negatively. In fact, they may help by making the microstructure denser. Less harmful pores range from 20 to 50 nm and can allow small amounts of water and gases to enter the concrete, which might cause minor damage under certain conditions. Pores ranging from 50 to 200 nm are considered harmful because they permit the greater penetration of water and detrimental chemicals into the concrete. This can lead to significant damage like freeze–thaw damage, reactions that weaken the concrete, or the corrosion of any metal inside it. Very harmful pores are larger than 200 nm and are the most problematic. They let aggressive agents enter freely, leading to severe deterioration and the reduction of the concrete’s lifespan, especially in harsh environments.
Table 6 presents the porosity for each pore size category in the SSAS with varying water-to-binder ratios. Initially, as the W/B ratio was raised from 0.43 to 0.51, the total porosity surged by 38.13%. This increase was reflected across all pore categories: harmless pores by 44.23%, slightly harmful pores by 177.42%, harmful pores by 22.22%, and very harmful pores by 8.6%. Such an increase in porosity across different categories significantly reduces the compactness of the structure, leading to a diminished compressive strength. Moreover, a looser internal structure not only facilitates easier CO2 penetration but also accelerates the migration of Cl− within the mortar, consequently lowering the carbonation resistance with the rising W/B ratio.
At a carbonation age of 28 days, the total porosity of the mortar rises along with the W/B ratio. When this ratio increases from 0.43 to 0.51, the total porosity goes up by 18.89%. The increases are distributed among the pore categories as follows: harmless pores by 9.87%, slightly harmful pores by 79.55%, harmful pores by 148.70%, and very harmful pores by 17.97%. This escalation in total porosity and category-specific porosity leads to a looser structure, which enables CO2 to diffuse more easily within the specimens. Consequently, as the W/B ratio climbs, both the carbonation resistance and the mechanical property of the SSAS decrease.
At a W/B ratio of 0.43, the total porosity of the mortar initially rises and subsequently declines with the progression of carbonation age. In the initial phases of carbonation, the decalcification of the C-(A)-S-H gel leads to a temporary increase in porosity. Subsequently, the pores begin to fill as calcium carbonate forms during the carbonation process. Despite this overall reduction in porosity, the porosity of the very harmful pores continues to increase, potentially accelerating the transport of Cl−, which could compromise the structural integrity of the mortar. Unlike smaller pores, the harmful pores (>200 nm) did not significantly decrease, which may be due to carbonation products like calcium carbonate preferentially filling the smaller pores [74]. This can lead to a denser microstructure, reducing the permeability to gases and liquids. However, the presence of larger pores can still negatively affect mechanical strength and durability, as they provide pathways for crack propagation and facilitate the ingress of harmful substances.
Figure 18 and Figure 19 illustrate the pore size distribution differential curves for the SSAS across various water–binder ratios and ages. Initially, under standard curing conditions before carbonation, the pore sizes for groups W0.43 and W0.51 are primarily concentrated in areas with sizes smaller than 100 nm. Most of the porosity in these mortar groups is composed of harmless pores. However, with an increase in the W/B ratio, there is a noticeable rise in both the overall porosity and the porosity at each specific pore size level. This increase is attributed to a reduction in the alkalinity of the mortar specimens, which impairs the activator’s capacity to enhance the slag reactivity, consequently slowing the hydration rate and reducing the production of hydration products, thereby increasing porosity [36]. A higher total porosity and a larger proportion of larger pores (especially harmful and more harmful pores) can significantly increase the permeability of concrete. This means that harmful substances like water, chloride ions, and carbon dioxide can more easily enter and penetrate the concrete structure, leading to an increased risk of corrosion for the internal reinforcement. In the pore structure, an increase in the proportion of harmless pores may have a dual effect on diffusivity. On one hand, small harmless pores can impede the rapid diffusion of some harmful substances; on the other hand, an increase in total porosity and the presence of medium to large pores can enhance overall diffusivity, allowing harmful substances to move more quickly within the concrete.
Following carbonation, the pore size distribution in the concrete broadens, and new peaks around 10,000 nm become apparent on the differential curve, signaling a worsening trend in pore sizes. The porosity in these groups shifts from predominantly harmless to a greater proportion of harmful pores. This transformation indicates substantial structural changes within the mortar, leading to a potential decrease in mechanical strength and durability due to heightened susceptibility to environmental penetrants such as CO2 and chlorides. After the carbonation process, the increase in the proportion of harmful and very harmful pores in the samples can lead to the enhanced diffusivity of the concrete. The presence of large and medium pores provides faster diffusion paths, enabling harmful substances to move more rapidly within the material. This accelerates the deterioration process of the concrete and shortens its service life.

4. Conclusions and Prospect

To explore the impact of the W/B ratio on the carbonation resistance of SSAS concrete and the distribution of chloride ions after carbonation, this study employed tests on the flowability, mechanical properties, carbonation performance, and chloride ion concentration of the mortar, alongside XRD, MIP, and TG-DTG microanalysis methods to investigate the underlying mechanisms. The primary conclusions are summarized as follows:
(1)
Mortar specimens with a lower W/B ratio demonstrate more complete hydration, resulting in lower porosity and enhanced mechanical properties, though they exhibit poorer flowability.
(2)
Increasing the W/B ratio results in a decrease in the quantity of hydration products, higher porosity, accelerated carbonation reaction rates, and faster chloride ion transmission and release.
(3)
During carbonation, the C-(A)-S-H gel in the samples initially undergoes decalcification, reducing the porosity and compressive strength. As carbonation continues, the generation of CaCO3 fills the pores, which decreases porosity and gradually enhances strength.
(4)
The carbonation process significantly influences the Cl− distribution. In the carbonated interface area of the mortar, chloride ion concentration is substantially higher than in non-carbonated areas. This indicates that carbonation drives the transmission of Cl− from the sea sand towards the carbonation interface, resulting in a peak concentration of chloride ions just ahead of this interface. Notably, samples from groups with higher water–binder ratios exhibit higher peak concentrations and deeper peak locations, suggesting a heightened risk of chloride ion erosion.
(5)
Our research on the interaction between carbonation and chloride ions in SSAS concrete has significant implications for developing predictive models and service life estimation tools. By incorporating these interactions into future models, we can more accurately predict the long-term durability of SSAS concrete structures under various environmental conditions.
(6)
The research findings of this study have multiple potential practical applications, including the design of more durable marine structures, the development of eco-friendly building materials, and the improvement of the corrosion resistance of reinforced concrete. Future research will focus on optimizing the mix design of SSAS concrete, exploring its long-term performance under various environmental conditions, and developing advanced models for predicting the service life of SSAS concrete structures. Further studies will also investigate the use of different types of sea sand and alkali activators to enhance the performance of SSAS concrete.

Author Contributions

Conceptualization, S.K. and J.Z.; Methodology, J.Z. and W.L.; Validation, Y.W., H.H. and W.L.; Formal analysis, Y.W., F.Z., J.Z. and W.W.; Investigation, Y.W.; Resources, S.K. and Z.Z.; Data curation, Y.W., S.K. and W.W.; Writing—original draft, S.K. and H.L. (Haojie Liu); Writing—review & editing, S.K., H.H. and H.L. (Haojie Liu); Visualization, Y.W. and Z.Z.; Supervision, F.Z. and H.L. (Hongze Li); Project administration, Y.W. and H.L. (Hongze Li); Funding acquisition, H.H. and W.W. All authors have read and agreed to the published version of the manuscript.

Funding

The authors express their gratitude for the financial support provided by the Xiamen municipal infrastructure project Tingxi West Road (2017-350212-78-01-000848), the Xiamen municipal infrastructure project Ring East Sea New City NO. 9 Binhai Road (2019-350212-78-01-60036), Fujian Provincial Department of Housing and Urban-Rural Development (2022-K-212), Natural Science Foundation of Fujian Province of China (2023J01998), and Engineering Research Center of Disaster Prevention and Mitigation of Southeast Coastal Engineering Structures of Fujian Province University (2022006).

Data Availability Statement

Data are contained within the article.

Conflicts of Interest

Authors Yan Wu, Jianbin Zhang and Zhou Zheng were employed by the company Xiamen Municipal Engineering Design Institute Co., Ltd. Author Feng Zhang was employed by the company CSCEC Strait Construction and Development Co., Ltd. Author Hongze Li was employed by the company Xiamen Municipal City Development and Construction Co., Ltd. Author Weihong Li was employed by the company Dalian Municipal Design and Institute Co., Ltd. The remaining authors declare that the research was conducted in the absence of any commercial or financial relationships that could be construed as a potential conflict of interest.

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Figure 1. Slag XRD diagram.
Figure 1. Slag XRD diagram.
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Figure 2. Ball mill front.
Figure 2. Ball mill front.
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Figure 3. Ball mill side.
Figure 3. Ball mill side.
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Figure 4. SSAS flowability for different W/B ratios.
Figure 4. SSAS flowability for different W/B ratios.
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Figure 5. Mechanical properties of SSAS at different W/B ratios.
Figure 5. Mechanical properties of SSAS at different W/B ratios.
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Figure 6. Carbonization depth test after carbonization of SSAS samples with different W/B ratios. (a) W0.43 carbonized for 3 days. (b) W0.43 carbonized for 7 days. (c) W0.43 carbonized for 14 days. (d) W0.43 carbonized for 28 days. (e) W0.47 carbonized for 3 days. (f) W0.47 carbonized for 7 days. (g) W0.47 carbonized for 14 days. (h) W0.47 carbonized for 28 days. (i) W0.51 carbonized for 3 days. (j) W0.51 carbonized for 7 days. (k) W0.51 carbonized for 14 days. (l) W0.51 carbonized for 28 days.
Figure 6. Carbonization depth test after carbonization of SSAS samples with different W/B ratios. (a) W0.43 carbonized for 3 days. (b) W0.43 carbonized for 7 days. (c) W0.43 carbonized for 14 days. (d) W0.43 carbonized for 28 days. (e) W0.47 carbonized for 3 days. (f) W0.47 carbonized for 7 days. (g) W0.47 carbonized for 14 days. (h) W0.47 carbonized for 28 days. (i) W0.51 carbonized for 3 days. (j) W0.51 carbonized for 7 days. (k) W0.51 carbonized for 14 days. (l) W0.51 carbonized for 28 days.
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Figure 7. Effect of W/B ratio on carbonation depth.
Figure 7. Effect of W/B ratio on carbonation depth.
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Figure 8. The SSAS compressive strength after carbonation with different W/B ratios.
Figure 8. The SSAS compressive strength after carbonation with different W/B ratios.
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Figure 9. Chloride ion concentration in specimens of different depths with different W/B ratios.
Figure 9. Chloride ion concentration in specimens of different depths with different W/B ratios.
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Figure 10. XRD of SSAS after 28 days standardized curing with different W/B ratios.
Figure 10. XRD of SSAS after 28 days standardized curing with different W/B ratios.
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Figure 11. XRD of SSAS after 28 days carbonated curing with different W/B ratios.
Figure 11. XRD of SSAS after 28 days carbonated curing with different W/B ratios.
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Figure 12. XRD of W0.43 SSAS as the carbonation progresses.
Figure 12. XRD of W0.43 SSAS as the carbonation progresses.
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Figure 13. TG-DTG of SSAS concrete with varying waterbinder ratios after 28 days standard curing.
Figure 13. TG-DTG of SSAS concrete with varying waterbinder ratios after 28 days standard curing.
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Figure 14. Weight loss of specimens at different temperature intervals after 28 days standard curing.
Figure 14. Weight loss of specimens at different temperature intervals after 28 days standard curing.
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Figure 15. TG-DTG of SSAS concrete with varying water–binder ratios after carbonation curing.
Figure 15. TG-DTG of SSAS concrete with varying water–binder ratios after carbonation curing.
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Figure 16. Weight loss of specimens at different temperature intervals after 28 days of carbonation curing.
Figure 16. Weight loss of specimens at different temperature intervals after 28 days of carbonation curing.
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Figure 17. The mass loss rate of SSAS mortar at different carbonation ages.
Figure 17. The mass loss rate of SSAS mortar at different carbonation ages.
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Figure 18. Differential curves of pore size distribution of SSAS with varying W/B ratios for 28 days of standard curing.
Figure 18. Differential curves of pore size distribution of SSAS with varying W/B ratios for 28 days of standard curing.
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Figure 19. Differential curves of pore size distribution of SSAS with varying W/B ratios after 28 days of carbonization.
Figure 19. Differential curves of pore size distribution of SSAS with varying W/B ratios after 28 days of carbonization.
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Table 1. Basic parameters of slag.
Table 1. Basic parameters of slag.
Density
(g/cm3)		Cl− Concentration (%)Fluidity
Ratio (%)		Loss on Ignition
(%)		Specific Surface Area (m2/kg)Activity Index of 7 Days (%)Activity Index of 28 Days (%)Water
Content (%)
2.920.0321031.0242578960.36
Table 2. The composition of slag/%.
Table 2. The composition of slag/%.
CaOSiO2Al2O3MgOSO3TiO2Na2OFe2O3MnOK2OClSrOOthers
36.8430.2717.189.672.791.250.600.480.400.340.050.040.10
Table 3. Specific substance content in RS and SS.
Table 3. Specific substance content in RS and SS.
RSSS
water content/%4.800.54
soil content/%2.461.52
Cl− content/%0.0030.163
Shell content/%/11.20
Table 4. Cumulative percentage of sieve residue for both sands.
Table 4. Cumulative percentage of sieve residue for both sands.
Cumulative Sieve Residue/%RSSS
A10.000.80
A20.2018.98
A35.2837.44
A480.8063.10
A597.7286.74
A699.4496.06
Table 5. Constituent parameters of SSAS.
Table 5. Constituent parameters of SSAS.
GroupB/S (Binder/Sand)W/B RatioSea Sand ContentAlkali DosageAlkali Modulus
W0.391:20.3930%6%1.0
W0.431:20.4330%6%1.0
W0.471:20.4730%6%1.0
W0.511:20.5130%6%1.0
Note: The sand used in this experiment is a mixture of sea sand and river sand. The sea sand content was 30% and the river sand content was 70%.
Table 6. Porosity of samples with varying carbonation age for different/%.
Table 6. Porosity of samples with varying carbonation age for different/%.
Group NumberCarbonization AgeTotal Porosity/%Harmless Pore (d ≤ 20 nm)Less Harmful Pore (20 < d ≤ 50 nm)Harmful Pore (50 < d ≤ 200 nm)More Harmful Pore (d > 200 nm)
W0.430d11.207.710.310.272.91
14d13.344.422.092.004.83
28d11.443.851.321.155.12
W0.510d15.4711.120.860.333.16
28d15.504.232.372.866.04
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MDPI and ACS Style

Wu, Y.; Kang, S.; Zhang, F.; Huang, H.; Liu, H.; Zhang, J.; Li, H.; Li, W.; Zheng, Z.; Wu, W. Study on the Effect of Water–Binder Ratio on the Carbonation Resistance of Raw Sea Sand Alkali-Activated Slag Concrete and the Distribution of Chloride Ions after Carbonation. Buildings 2024, 14, 2027. https://doi.org/10.3390/buildings14072027

AMA Style

Wu Y, Kang S, Zhang F, Huang H, Liu H, Zhang J, Li H, Li W, Zheng Z, Wu W. Study on the Effect of Water–Binder Ratio on the Carbonation Resistance of Raw Sea Sand Alkali-Activated Slag Concrete and the Distribution of Chloride Ions after Carbonation. Buildings. 2024; 14(7):2027. https://doi.org/10.3390/buildings14072027

Chicago/Turabian Style

Wu, Yan, Sixiang Kang, Feng Zhang, Haisheng Huang, Haojie Liu, Jianbin Zhang, Hongze Li, Weihong Li, Zhou Zheng, and Wenda Wu. 2024. "Study on the Effect of Water–Binder Ratio on the Carbonation Resistance of Raw Sea Sand Alkali-Activated Slag Concrete and the Distribution of Chloride Ions after Carbonation" Buildings 14, no. 7: 2027. https://doi.org/10.3390/buildings14072027

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