Study on Load-Bearing Characteristics and Engineering Applications for Cement–Soil Pipe Pile

Abstract

The cement–soil pipe pile is a novel blend of cement and soil, designed to enhance load-bearing capabilities while cutting down on the need for cement. Its tubular construction is key to its strength. To delve into how the pile’s cross-sectional size affects its load-bearing properties, we took into account the soil–cement’s strain-softening behavior. Laboratory tests examined the load-bearing properties of piles. We created an exponential decay Mohr–Coulomb model in ABAQUS for further development, performed field tests, and built a numerical model incorporating wall thickness, pile diameter, and length. The unit volume ultimate bearing capacity was used to evaluate pile performance, with a focus on a 600 mm diameter pile. The results show that wall thickness minimally affects load-bearing capacity, needing to be at least a quarter of the diameter. Larger diameters increase the ultimate bearing capacity, but the capacity per unit volume declines. The 600 mm diameter pile boasts the highest unit volume ultimate bearing capacity. The pile’s effective length is roughly 10 m. Beyond this, extending the pile length increases the single pile’s ultimate bearing capacity by less than 5%, but the unit volume capacity starts to dwindle.

1. Introduction

As an economical, convenient, and effective measure for soft ground reinforcement treatment, cement–soil mixing piles have been widely used in highway engineering [1], railway engineering [2], support engineering [3], offshore engineering [4,5], and hydraulic structure engineering [6]. Hydraulic soil mixing piles have really taken off in highway engineering because they are a relatively cheap option. That being said, it is not all sunshine and roses. One major snag is that the soil’s, shall we say, personality, layer by layer, can really throw a wrench into the pile formation process. And if the quality takes a nosedive, the piles might not be able to shoulder the load they are meant to. So, there is a real and pressing need to come up with a new and improved way of building cement–soil piles that gives us tighter control over quality as we are building, gets the most bang for our buck in terms of materials, and gives those piles a serious boost in load-bearing capacity.

Cement–soil pipe piles represent an innovative approach to composite foundation treatment, building upon traditional cement–soil mixing piles to enhance soft ground stability. This technique offers a range of advantages, such as a higher load-bearing capacity, optimized material usage, and enhanced construction efficiency. The formation of cement–soil pipe piles involves a tubular design that is an improvement over traditional mixing drill bits (refer to Figure 1). Specifically, this method requires an increase in the slurry spraying pressure and a modification of the spray nozzle’s position. Consequently, the soil within the pile’s core remains intact, while a layer of cement slurry coats the soil core, running from the spray nozzle down to the length of the mixing rod, thereby creating a cylindrical cement–soil pipe pile. Utilizing the same quantity of materials, this approach permits a larger pile diameter and improved load-bearing capability. By boosting the bearing capacity of each individual pile, we can also widen the spacing between them, which in turn diminishes the overall number of piles needed. To address the problem of the poor pile formation quality of soil-mixed piles, Liu Songyu et al. [7,8,9,10] investigated the ultimate single-pile bearing capacity, damage model, and optimum shape design of expanding head of nail-type bi-directional mixing piles by means of indoor tests, field tests, and numerical simulations. However, considering the differences in the structural forms of the above solid-core piles and cement–soil pipe piles, the bearing mechanism is obviously different from that of cement–soil pipe piles. The large-diameter cast-in-place concrete pipe piles have a high similarity with cement–soil pipe piles in terms of pile structure. The former is a rigid pile, and the latter is a flexible pile. Liu Hanlong et al. [11,12,13,14] have systematically elaborated the development, design, and field application of this technology, which is a useful reference for the study of this paper. Wang Zhe and colleagues [15] investigated how the soil core within large-diameter cast-in-place concrete pipe piles affects the load-bearing capacity of the piles themselves. They suggested that the section of the soil core contributing to inner friction resistance is roughly one-third of the pile’s length from the bottom and that the bearing capacity derived from the soil core represents approximately 7% of the overall load-bearing capacity. J Xu [16,17] considered the influence of different model parameters on relative displacement and stress, which is a useful reference for the study of this paper. Cement–soil pipe piles improve soft ground reinforcement through a tubular design, enhancing pile quality and load capacity while reducing cement and pile quantity, yielding economic benefits. Addressing the insufficient understanding of hydraulic clay’s strain-softening in bearing capacity [18,19], this study combines in situ testing and numerical modeling to analyze the impact of pipe pile dimensions on load-bearing capabilities. The results will inform practical engineering applications, providing both theoretical and practical guidance.

2. Field Experiment

The field application of cement–soil pipe piles was carried out in conjunction with a specific project in Shandong Province, and a field test pile was carried out prior to the project application. The test pile site was the open space on the inner side of the ramp where the project application was to be carried out. The overburdened soil layer at the site was mainly powdery clay, chalk, and silt, which is Quaternary Holocene (Q4).

The mixer pile driver was ZGZ-B type with a main motor power of 55 kW; the mud pump was ZB2-100 type. The motor power was 15 kW, the rated pressure was 20 Mpa, and the flow rate was 70 L/min. The construction process of cement–soil pipe pile was as follows:

(1)

The pipe pile should adopt a special pipe pile churning and spraying drill bit, and it is preferable to adopt the sinking drilling and spraying method as a priority.

(2)

When the driving pipe pile spraying and stirring bit sinks into the ground, start the high-pressure mud pump to supply cement slurry to the nozzle, and turn off the mud pump while drilling and spraying and stirring until the designed depth is reached; the value of the cement slurry spraying pressure is 10 Mpa–12 Mpa (adjusted according to the wall thickness).

(3)

Stirring and lifting at the same time, turn on the mud pump for re-spraying and re-stirring when it reaches a height of 2.5 m from the top of the pile. Stop the slurry after the drill is lifted to the ground elevation.

Improvements can be made to traditional cement–soil pipe pile construction equipment to enhance its effectiveness. First, it is essential to reposition the drill bit nozzle and modify its dimensions so that the distance from the nozzle to the end of the stirring rod matches the thickness of the pile wall (refer to Figure 1). Additionally, the high-pressure nozzle should be adjusted to a size between 1.5 and 2.2 mm to maintain optimal slurry spraying pressure. A high-pressure pump should be employed for the slurry application. The construction method consists of two stirring actions followed by one spraying action: While the pile driver’s drill bit penetrates the ground, the slurry is both sprayed and stirred; when the drill bit is being retracted, only stirring occurs. The drilling lift speed should be regulated between 45 and 80 cm/min, depending on the soil layer characteristics, while the spraying pressure needs to be maintained at 10–12 Mpa. Throughout the construction process, it is crucial to monitor any fluctuations in spraying pressure and ensure the drilling rod’s vertical alignment. Work should be halted if any irregularities arise, allowing time to diagnose issues, rectify faults, and adjust parameters before resuming construction. If the spraying and stirring processes need to be temporarily paused, it is important to restart the spraying with an overlap of at least 0.3 m. Keeping detailed construction records is vital for maintaining quality standards throughout the project.

Two groups of Φ600 mm and Φ800 mm diameters were selected for the in situ test, and each group of two piles was formed into three piles with a length of 10 m. The static load test of single pile bearing capacity and core drilling sampling were carried out. The on-site testing method refers to the “Code for the Design of Building Foundations” (GB50007-2011). The field test pile excavation and core drilling test are shown in Figure 2 and Figure 3, respectively, and the results of single pile ultimate bearing capacity test are shown in Figure 4.

Twenty-eight days following the conclusion of the testing, the in situ test piles on-site underwent evaluation through pile head excavation and core drilling sampling. As illustrated in Figure 2, these in situ test piles yielded promising results, thanks to an enhanced construction methodology that resulted in hydraulic clay piles shaped in a tubular form. The core samples retrieved from the drilling exhibited completeness, with the majority being columnar in shape and a minor portion appearing lumpy, all while maintaining a uniform mixture. The anticipated objectives were successfully met. Figure 4 presents the displacement curves for the bearing capacity of single piles with diameters of 600 mm and 800 mm. Notably, the ultimate bearing capacity for the 600 mm tubular piles stands at 430 kN, while the 800 mm counterparts reach 785 kN, representing increases of 53% and 180%, respectively, over the original design requirement of 280 kN. It is important to highlight that the cement material quantity for the 600 mm piles matches that of the originally designed 500 mm solid piles, which is set at 60 kg/m. In contrast, the cement requirement for the 800 mm piles is only 92% higher than the original design, with the former requiring 60 kg/m and the latter 115 kg/m.

3. Physical Model Testing

Combined with the actual engineering situation, simplifying the related engineering into a similar model with similar materials, and carrying out targeted multi-parameter, multi-angle, and multi-condition analysis, can solve the disadvantages of irreversible and susceptible environmental impact of field tests. Based on the similar model test, the key parameter information can be obtained in an all-round and three-dimensional way by embedding relevant monitoring elements in advance, and the bearing characteristics of the pile foundation and the deformation characteristics of the soil can be intuitively and vividly revealed.

3.1. Model Testing System

The model test system determined by the bearing characteristic test consists of three parts: the model tank, loading system, and measurement system. Among them, the loading system is used for the bearing characteristic test research of the stirred pile, the measurement system is mainly based on the bearing capacity and stress-strain testing technology, real-time measurements of the pile force, soil deformation, and strain are carried out in the test process, the model tank is the main part of the model test to carry out and implement the model test, and it is possible to simulate different test sites by changing the filling materials in the tank. The indoor model test system is shown in Figure 5. In the field test, the length of the mixing pile is 10 m; in order to increase the operability of the model test, the volume of the test model cannot be too large, and at the same time, in order to ensure the accuracy of the test, the volume of the test model cannot be too small. According to the literature and on-site construction experience, the geometric similarity ratio was C_l = 15, the severe similarity ratio was C_γ = 1, the length of the model pile was 66.7 cm, and the pile diameter was 40 mm. The model box was 1.5 m in length, width, and height.

3.2. Test Conditions

The load-bearing characteristic test conditions are shown in

Table 1. Test program for single pile bearing characteristics.

Working Condition	Serial Number	OD/mm	ID/mm	Pile Length/mm	Note
PS	PS	33	33	667
	PT-1	40	20	667	effect of cylinder pile size
PT	PT-2	47	23	667
	PT-3	53	27	667

Note: PS—plain solid core pile; PT—equal cross-section cylindrical pile.

.

3.3. Model Building

(1) Model pile making

The model pile is made by 3D printing, after research; in many materials, white resin has stable performance and is easy to mold, other materials are too hard and close to the rigid body, and white resin is relatively soft, which can better simulate the bearing characteristics of the hydraulic soil cylinder pile, so it is determined that the material is white resin, which simulates the bearing characteristics of the hydraulic soil cylinder pile. The product parameters are shown in

Table 2. Physical properties—liquid material.

Appearance	White
Density	1.11~1.15 g/cm3 @25 °C
Viscosity	280~420 cps @ 25 °C
Dp	0.135~0.158 mm
Ec	8.1~9.0 mJ/cm2
Building 1ayer thickness	0.05~0.12 mm

and

Table 3. Mechanical Properties of Post-Cured Material.

Measurement	Test Method	Value
		90 min UV post-cure
Hardness Shore D	ASTM D 2240	76~88
Flexural modulus Mpa	ASTM D 790	2692~2775
Flexural strength Mpa	ASTM D 790	69~74
Tensile modulus MPa	ASTM D 638	2589~2695
Tensile strength MPa	ASTM D 638	38~56
Elongation at break	ASTM D 638	8~12%
Poisson’s ratio	ASTM D 638	0.4~0.44
Impact strength notched Izod, J/m	ASTM D 256	32~38
Heat deflection temperature, °C	ASTM D 648@66 PSI	39~52
Glass transition, Tg °C	DMA, E” peak	40~57
Coefficient of thermal expansion/°C	TMA(T)	90~103 × 106
Density g/cm3		1.12~1.18
Dielectric constant 60 Hz	ASTM D 150-98	4.2~5.0
Dielectric constant 1 kHz	ASTM D 150-98	3.3~4.2
Dielectric constant 1 MHz	ASTM D 150-98	3.2~4.0
Dielectric strength kV/mm	ASTM D 1549-97a	12.8~16.1

.

(2) Test fill

The source of soil used in this test was standard sand. The model arrangement of the single pile bearing characteristic test is shown in Figure 6.

3.4. Test Methods

The interior of the model trench was filled with on-site soil in layers from the bottom to the top, and every 10 cm of fill was rolled in layers and compacted according to the similarity coefficient of dry density C_γ. After filling to the design depth, model piles were installed. One strain gauge was set at 10 cm intervals on the pile side of the model pile, and a total of seven sets of strain gages were set.

The loading test was carried out using the loading system, which employed a displacement control method to apply the vertical load. The entire setup is secured to the sliding rail of the model tank via a crossbeam and a detachable bracket. At the front end of the hydraulic system, a spoke-type pressure transducer and a micrometer were installed. Throughout the loading test, strain gauges and pressure transducers were linked to the data acquisition system to track the variations in load, displacement, and stress parameters during the loading phase.

3.5. Test Steps

Before the start of the test, the bottom of the test should be filled in advance and reserve a good space for burying the test pile and the prepared fill soil uniformly and symmetrically filled to the model box and layered paving, Model box pre-filled soil as shown in Figure 7.

Currently, there are two ways of burying model piles, i.e., the press-in type and the buried type. The press-in type is generally used to simulate soil crowding piles; the specific method is to first fill the soil to a predetermined pile top height position and then press the model pile into the soil. The buried type is mainly used to simulate non-extruded piles; the specific method is to first fill the soil to a predetermined height of the pile end, place the model pile in the specified position and fix it, and then continue to fill the soil to a predetermined height of the top of the pile. The buried type was used in this test.

Positioning the model stakes is a very important step in the process of burying the model stakes. Before burying the model pile, first set out the tempered glass around the model box and the centerline of the model pile. In order to ensure the accurate positioning of the model piles, the centerline of the tempered glass and the centerline of the model piles should be made to coincide with each other in the process of burying the model piles.

Soil filling in the model box must be performed with care to minimize the effect of soil filling on the positioning of the model piles. Complete the preparatory work of the model pile as shown in Figure 8.

The data measurement system of this model test include the following two aspects: pile top settlement and pile stress measurement. Among them, the settlement of the pile top is measured by the percentage meter arranged at the top of the pile, as shown in Figure 9. And the strain measurement of the pile body adopts the DH3618N static strain testing system produced by Donghua Testing Co. (Suzhou, China).

Vertical loading tests were performed on the model piles after all preparations were made. The status of all work completion is shown in Figure 10.

The specific operation steps are as follows:

This model test was loaded step by step during the test, 200 N for each level, and then the next level of load was added after each level of load reached relative stability, until a certain level of load of the model pile reached twice the settlement of the previous level of load. Finally, according to the termination loading condition of Article 2 of 6.3.4 in the “Technical Code for Building Foundation Testing” JGJ 340-2015 (under a certain level of load, the settlement of the pile top is greater than 2 times the settlement of the previous level of load), the ultimate bearing force standard was set in the laboratory test.

(1)

The control model pile is loaded with 200 N until the ultimate load is obtained.

(2)

The settlement observation specification stipulates that the next level of loading shall not be carried out when the subsidence is not stabilized. The observation time of each level of loading is as follows: Each level of loading should be observed immediately after loading, and then in the first hour, it is observed every 15 min; in the second hour, it is observed every 30 min; from the third hour onwards, it is observed every 1 h.

(3)

Stabilization standard of settlement: If the amount of subsidence of each level of load is not more than 0.1 mm in the last 30 min, the settlement can be regarded as stabilized, and the next level of load can be applied.

(4)

The termination of loading and the value of limit load: The loading can be terminated when the sinking amount of the current load is equal to or greater than 2 times the sinking amount of the previous load.

The specific case of loading is shown in Figure 11.

3.6. Experimental Results and Analysis

For this experiment, four identical piles were constructed, each divided into equal sections. The displacement of the pile’s apex was monitored using a percentage meter installed at the top, while the load pressure was regulated with a pressure transducer.

Analyzing the test outcomes for each model pile allows for a comparison of the pile top load–settlement curves (P-S curves), as illustrated in Figure 12.

From the “Technical Specification for Pile Foundation of Construction”, it is reasonable to consider the uppermost layer of the preceding stack, where the settlement induced by a specific load is double that of the load before it, as the ultimate bearing capacity for the model pile. From this, we can conclude that the vertical ultimate bearing capacity of the PS model pile (a uniformly solid pile) is 1000 N, the PT-1 model pile (equal-section cylindrical pile) is 1400 N, the PT-2 model pile (equal-section cylindrical pile) is 1800 N, the PT-3 model pile (equal-section cylindrical pile) is 2400 N, and the ST-1 model pile (variable-section cylindrical pile) is 2400 N.

According to the strain results of this indoor model test, the distribution law of pile body axial force can be analyzed. The strain value at the corresponding position of the pile body is measured by the resistance strain gauges attached to the pile body at the corresponding position, and then the axial force at the corresponding cross-section position can be obtained by the stress–strain conversion formula in the mechanics of materials:

$$Q_i=EA{\epsilon }_i$$

(1)

Type: Q_i is the pile body axial force (N); E is the modulus of elasticity of the pile body (2.1 GPa); A is the cross-sectional area of the pile body (m²); and ε_i is the axial strain at pile body I (με).

The axial force data and the distribution of the axial force of each model pile under different loading levels can be obtained from the conversion of the strain data of the pile.

As illustrated in Figure 13, there is a noticeable trend: As the load on the pile top rises, the relative movement between the pile and the surrounding soil also steadily increases. Additionally, the axial force experiences a gradual uptick, with the upper section of the pile showing a significantly quicker rise in axial force compared to the lower section. This indicates that the axial force within the pile is progressively being transmitted from the top toward the bottom, ultimately reaching the pile’s tip.

As illustrated in Figure 14, both the ultimate bearing capacity of a single pile and the unit volume bearing capacity of equal-section cylindrical piles surpass those of solid piles. Furthermore, it is evident that the ultimate bearing capacity of these cylindrical piles tends to increase with a larger outer diameter. This enhancement occurs because a greater outer diameter expands the contact area between the pile and the surrounding soil, both on the sides and at the pile’s base, with a notable increase on the lateral surfaces. However, calculations reveal that a larger outer diameter does not always equate to better performance. Specifically, when evaluating cylinder piles with outer diameters of 40 mm, 47 mm, and 53 mm, the respective ultimate bearing capacities per unit volume were found to be 2.23 N, 2.05 N, and 2.02 N. Notably, the 53 mm diameter pile exhibited a reduction of approximately 10% in ultimate bearing capacity per unit volume compared to the 40 mm pile. Therefore, it is crucial to judiciously select the outer diameter of cylindrical piles in practical applications.

4. Computational Modeling

4.1. Mohr–Coulomb Model for Exponential Decay of Hydraulic Soils

Conventional soil constitutive models are unable to simulate the strain-softening characteristics of hydraulic soils under loading [18,20]. As illustrated in Figure 15, Yapage et al. [20,21] put forth a single-fold strain softening model. In this model, strain softening initiates when the plastic shear strain hits the critical threshold, denoted as ε_d0. At this point, the material’s microelemental strengths—specifically the internal friction angle and cohesion—exhibit a linear decline towards their residual values, while the shear expansion angle tapers off to zero in a linear fashion. Although this approach is straightforward and effective, it falls short since the reduction of hydraulic soil strength parameters is not linear [22,23]. To more accurately capture the strain-softening behavior of hydraulic soil, an exponential decay model was proposed, as presented in Equation (2):

$$C=C_0{e}^{-\left(\epsilon -{\epsilon }_{d0}\right)D_c}+(1-e^{-\left(\epsilon -{\epsilon }_{d0}\right)D_c})C_r$$

(2)

Calculation formula: C₀ is the initial value of microelement strength, such as cohesion, the friction angle, etc.; ε_d0 is the softening onset critical plastic shear strain; C_r is the residual value of microelement strength after softening; and D_C is the decay parameter.

4.2. Triaxial Test Validation

In order to verify the accuracy of the above exponential decay model, numerical simulations were carried out based on triaxial tests in the literature [20,21]. The model is axisymmetric, and the material parameters are consistent with the literature (Young’s modulus E, peak cohesion c, post-peak cohesion C_r, peak friction angle φ, post-peak friction angle φr, and softening onset critical plastic shear strain ε_d0). The mesh is a four-node axisymmetric quadrilateral cell (CAX4P).

Figure 16 presents the simulation results alongside findings from the literature. It is evident from this figure that the conventional Mohr–Coulomb model falls short in capturing the strain-softening behavior of hydraulic soil. Initially, when loading begins and the hydraulic soil has yet to enter the softening phase, both the single-fold strain softening model and the exponential decay model exhibit similar stress–strain behavior. However, as loading progresses and the soil transitions into the softening stage, the exponential decay model aligns more closely with the deformation patterns observed in hydraulic soil, thereby providing a more accurate representation of its strain softening characteristics compared to the single-fold strain softening model.

4.3. Field Test Validation

The numerical analysis presented in this paper is grounded in the field test outlined in Part 2, utilizing the exponential decay Mohr–Coulomb model as previously described and executed using the ABAQUS 2022 finite element software. To enhance the accuracy of the model outcomes and address the impact of boundary effects, as well as the challenges posed by axisymmetric vertical forces acting on the pile, we have made sure the soil’s width is at least 20 times broader than the pile’s size. Moreover, the gap between the pile’s top and the soil’s bottom is usually maintained at more than 10 times the pile’s width. A two-dimensional axisymmetric numerical model has been created, illustrated in Figure 17. The dimensions of the foundation soil are 20 m (33 d) by 40 m (66 d), with the soil layers in the model accurately reflecting actual depths while being reasonably simplified. The pile dimensions correspond to those obtained from the in situ testing.

In this model, we establish the contact interface between hydraulic soil and regular soil. The tangential friction at this interface is characterized by a “penalty” method, with a friction coefficient defined as tanφ = 0.45, where φ represents the internal friction angle [24,25]. The contact surface is treated as hard contact in the normal direction, which effectively prevents penetration during the calculations. The contact pair is defined by stiffness, with the hydraulic soil designated as the master surface due to its higher stiffness, while the soil body functions as the slave surface [24,25]. The soil body itself is modeled using the Mohr–Coulomb approach. Displacements in the X and Y directions at the model’s bottom boundary are fixed at zero, as are the X direction along the pile axis and the left and right boundaries of the soil model. The loading strategy follows an incremental force control method based on field tests, with initial geostatic stress settings and equilibrium adjustments in place. The model employs a four-node axisymmetric cell, referred to as CAX4.

The parameters of each soil layer and pile body in the calculation model are selected according to the parameters determined by the field geological investigation data and related literature [26,27,28,29], and the summary values of the parameters of the calculation model are shown in

Table 4. The parameters of numerical simulation model for soil.

Profundity/m	Material Type	Severe/ (kN/m3)	Cohesion/ kPa	Internal Friction Angle/°	Young’s Modulus/ MPa	Poisson’s Ratio
0–5	Silty Clay	17	32	28	24.5	0.3
5–10	Dust	18	20	20.5	22.5	0.3
10–40	Silt	18.5	12	28	17.5	0.3

.

The ratio of C_r/C₀ in Equation (2) is defined as the residual softening index. Based on the research of Yapage et al. [20,21], and combined with field tests, the residual softening index is taken as 0.6 in the numerical simulation of this paper. According to the research of scholars, such as Navin and Filz [30,31,32], the cement–soil pipe pile is taken to be 100–300 times its unconfined compressive strength, which is taken to be 225 MPa (150 q_u, indoor measured q_u = 1.5 MPa). The internal friction angle of cement soil is generally 25~45° [33,34], which is taken as 35° in this paper, the cohesion cp is obtained with Equation (3) [33,34], and the material parameters for the numerical simulation of the cement–soil pipe pile are shown in

Table 5. The parameters of numerical simulation model for cement–soil pipe pile.

Modulus of Elasticity (kPa)	Cohesion (kPa)	Angle of Internal Friction (°)	Residual Value of Cohesion (kPa)	Residual Value of Angle of Internal Friction (°)	Cr/C0	Critical Plastic Shear Strain εd0 (%)
225 × 103	412	35	164.8	21	0.6	1

.

$$c_p={\displaystyle \frac{q_u}{2\tan \left(45+{\varphi }_p/2\right)}}$$

(3)

In the calculation equation, cp is the cohesive force (kPa) of the hydraulic soil, and φ_p is the angle of internal friction of the hydraulic soil.

Table 6. Numerical simulation and field measurement value comparison table.

Φ600 mm Single Pile Bearing Capacity (kN)			Φ800 mm Single Pile Bearing Capacity (kN)
On-Site Measurement	Exponential Decay Mohr–Coulomb	Mohr–Coulomb	On-Site Measurement	Exponential Decay Mohr–Coulomb	Mohr–Coulomb
430	445	445	785	670	670

presents the numerical simulation findings regarding the ultimate load capacity of individual piles. Initially, the Q-S behaviors for the Φ600 mm and Φ800 mm cement–soil piles, derived from both the exponentially decaying Mohr–Coulomb model and the ideal Mohr–Coulomb model, are quite similar during the early loading phase. The ultimate bearing capacities identified for the single piles are 445 kN and 670 kN, respectively. This similarity is attributed to the fact that, at this stage, the cement–soil piles have yet to enter the soft decay phase, as illustrated by the consistent results from both models. As vertical loading continues to rise, the Φ600 mm pile begins to soften once the external load surpasses 563.2 kN, while the Φ800 mm pile enters the softening phase when the load exceeds 760 kN. Notably, the exponential decay Mohr–Coulomb model, which accounts for diminishing internal friction angles and cohesive forces in response to increasing external loads, proves to be more representative of real-world engineering scenarios compared to the traditional Mohr–Coulomb model.

4.4. Calculate the Operating Conditions

The cross-sectional dimensions of a cement–soil pipe pile represent a novel structural form that greatly affects its load-bearing capacity. To explore how varying cross-section sizes impact the pile’s performance, a comparative analysis was conducted. This study examined the ultimate bearing capacity of the pile under different lengths, diameters, and wall thicknesses, drawing on data obtained from in situ tests conducted at the site. A summary of the calculated working conditions is presented in

Table 7. The conditions of numerical simulation.

Operating Condition	Outer Diameter/mm	Inner Diameter/mm	Pile Length/m	Pipe Wall Thickness/mm
1	600	0	10	300 (Solid Core Pile)
2	600	300	10	150
3	600	200	10	200
4	600	100	10	250
5	700	350	10	175
6	800	400	10	200
7	1000	500	10	250
8	600	300	6	150
9	600	300	8	150
10	600	300	12	150

.

5. Analysis of Calculation Results

In order to enable a more accurate comparison and analysis of the load-bearing capabilities of different cement–soil pipe pile sections, we propose the idea of the ultimate bearing capacity per unit volume. This concept is calculated by taking the ultimate bearing capacity of an individual pile and dividing it by the pile’s total volume. Essentially, this gives us a way to gauge the ultimate bearing capacity relative to the volume of the pipe pile itself. This index acts as a benchmark for assessing the load-bearing efficiency of each cement–soil pipe pile. The resulting variation curves, which depict the correlation between a single pile’s ultimate bearing capacity and its ultimate bearing capacity per unit volume, are illustrated in Figure 18, Figure 19 and Figure 20, showcasing various soil core diameters, pile diameters, and lengths.

5.1. Analysis of the Influence of Cement Soil Pipe Pile Wall Thickness on Bearing Capacity

In order to perform a comparative study on the impact of wall thickness on the load-bearing capacity of piles, a diameter of 600 mm was selected for the piles, with wall thicknesses ranging from 150 mm to 300 mm at increments of 50 mm for the solid core variants. As depicted in Figure 18a, the ultimate bearing capacity of the individual pipe piles varies with the different wall thicknesses. The findings indicate a steady increase in the ultimate bearing capacity as the wall thickness rises, with recorded values of 445 kN, 465 kN, 475 kN, and 476 kN, respectively. However, the rate of increase diminishes with each increment. Additionally, while thinner walls result in lower bearing capacity, they also require less cement material. To enhance our comprehension of the load-bearing characteristics of pipe pile structures, a comparative study was undertaken to examine how variations in wall thickness affect both the ultimate bearing capacity of individual piles and the ultimate bearing capacity per unit volume. As depicted in Figure 18b, the ultimate load-bearing capacity of a single pile shows a steady increase with a greater wall thickness. In contrast, the ultimate load-bearing capacity per unit volume decreases consistently, with measurements of 209 kN/m³, 185 kN/m³, 172 kN/m³, and 168 kN/m³ for wall thicknesses between 250 mm and 300 mm. Notably, the reduction in the ultimate load-bearing capacity per unit volume remains relatively insignificant within this range. However, when the wall thickness is increased from 150 mm to 250 mm, the ultimate capacity per unit volume decreases significantly by about 17%. Considering the low strength of hydraulic clay material, unlike cement concrete, the thin wall of a tubular pile easily causes damage to the hydraulic clay pile body [35]; therefore, the minimum thickness of tubular wall selected in this paper is 1/4 of the pile diameter.

5.2. Analysis of the Effect of Pile Diameter on Bearing Capacity

To investigate the impact of varying pile diameters on their bearing properties, calculations were performed to determine the ultimate bearing capacity of single piles with diameters ranging from Φ600 mm to Φ1000 mm. The internal core diameter was kept between 300 mm and 500 mm, while the wall thickness was maintained at 1/4 of the diameter (D). As depicted in Figure 19a, a noticeable trend emerges: The ultimate bearing capacity of a single pile increases with a larger diameter. Specifically, the recorded ultimate bearing capacities for diameters of Φ600 mm, Φ700 mm, Φ800 mm, and Φ1000 mm are 445 kN, 562 kN, 686 kN, and 986 kN, respectively. Furthermore, Figure 19b illustrates the connection between the ultimate bearing capacity of a single pile and the ultimate bearing capacity per unit volume as the diameter increases. Although the ultimate bearing capacity rises with increasing diameter, so does the overall volume of the pile and the cement required. As a result, the ultimate bearing capacity per unit volume decreases with larger diameters; notably, as the pile diameter grows from 600 mm to 1000 mm, the values drop to 209 kN/m³, 194 kN/m³, 182 kN/m³, and 161 kN/m³, indicating an approximate 23% decrease at the maximum. Therefore, from a cost-effective standpoint, opting for a cement–soil pipe pile with a diameter of Φ600 mm emerges as the most advantageous option, offering the greatest bearing capacity contribution for each unit of cement consumed.

5.3. Analysis of the Effect of Pipe Pile Length on Bearing Capacity

Figure 20 presents a comparison and analysis of how varying the lengths of tubular piles affect their load-bearing capabilities. As illustrated in Figure 20a, extending the pile length from 6 m to 12 m results in a rise in the ultimate bearing capacity of an individual pile, escalating from 336 kN to 407 kN, then to 445 kN, and finally reaching 456 kN. This translates to increases of 21%, 32%, and 35%, respectively. The effect of bearing capacity improvement is gradually smaller with the increasing length of the pile. This result indicates that for this study, an effective pile length also exists for cement–soil pipe piles, which are around 10 m long. It is consistent with the results of existing studies on effective pile length under embankment loading [36]. Figure 20b outlines the evolving trends in the ultimate load-bearing capacity of an individual pile and the ultimate load-bearing capacity per cubic unit of pipe pile, taking into account various lengths. As the length of the pile grows, its ultimate load-bearing capacity rises, while its capacity per unit volume decreases. This suggests that the unit mass of cement contributes less to the overall load-bearing capacity. Moreover, the effective pile length is influenced by a multitude of factors, such as the soil characteristics surrounding the pile, the composition of the pile material, and the loading circumstances. It is not a static value [37].

6. Engineering Applications

To further validate the use of cement–soil pipe piles, we built upon prior research findings and took into account the specifics of the current project. We opted for 600 mm diameter cement–soil pipe piles, aligning their length with the original design parameters—long piles measuring 11 m and short piles at 7 m. The site’s geological conditions were outlined in section one. The characteristic single-pile bearing capacity for the 600 mm piles was estimated at 215 kN. The spacing between the piles was determined using the replacement ratio “m” derived from the composite foundation bearing capacity calculation formula, with the final results presented in

Table 8. Comparison between the original design scheme and present design scheme.

Zones	Pile Spacing/m		Eigenvalue of Bearing Capacity of Single Pile/kN		Eigenvalues of Bearing Capacity of Composite Foundations/kPa	Number of Piles/m		Cement Usage/t
	Devise	Pipe Pile	Devise	Pipe Pile		Devise	Pipe Pile	Devise	Pipe Pile
A	1.2	1.5	140	213	180	11256	7475	617	447
B	1.5	1.8	140	213	145
C	1.8	2.2	140	213	135

. This replacement ratio allowed us to calculate the spacing for the pipe piles, which is also detailed in Table 8. The original design indicated pile spacings of 1.2 m, 1.5 m, and 1.8 m, whereas the new cement–soil pipe piles were spaced at 1.5 m, 1.8 m, and 2.2 m. Consequently, the total length of the pile extensions decreased from 11,256 m in the original design to 7475 m, a reduction of 3781 m, which represents over a 33% decrease. Additionally, the quantity of cement utilized dropped from 617 tons to 447 tons, a savings of 170 tons or more than 27%. In comparison to the initial design, the overall project costs have been lowered by over 25%, yielding significant economic and social advantages, along with considerable potential for future application and promotion.

To validate the efficacy of the tubular pile design, a series of field tests were carried out to evaluate the load-bearing capabilities of both individual piles and integrated foundations. The outcomes of these evaluations are depicted in Figure 21 and Figure 22. The data suggest that the load-bearing capacities of both the single piles and the composite foundations within three distinct regions meet the established design criteria. Specifically, the settlement measurements for the single pile bearing capacity were recorded as 14.92 mm in Zone A, 2.66 mm in Zone B, and 2.22 mm in Zone C, yielding an average settlement of 6.60 mm. For the composite foundation bearing capacity, the settlement values were 9.35 mm in Zone A and 4.59 mm in both Zone B and Zone C, leading to an average settlement of 6.43 mm. Notably, the settlement values associated with the 600 mm cement–soil pipe piles were lower, indicating a more effective foundation treatment.

7. Conclusions and Prospects

7.1. Conclusions

Field in situ test research and numerical simulation calculations have been carried out for the self-developed new cross-sectional form of cement–soil pipe piles to analyze the influence of the change in the cross-sectional dimensions of the pipe piles on the bearing characteristics of the pile body and to further carry out on-site engineering validation, and the following conclusions have been drawn:

• Drilling head and construction techniques for cement–soil pipe piles were developed, their reliability through on-site testing was validated, and a novel method for composite foundation treatment was proposed.
• As the diameter of equal cross-section cylinder piles increases, their single-pile bearing capacity rises, while the unit volume bearing capacity initially increases and then decreases, with the axial force showing a pattern of larger results upwards and smaller downwards.
• An exponential decay model using plastic shear strain was introduced, where cohesion, friction angle, and shear expansion angle decrease exponentially with shear plastic strain, effectively modeling strain softening in cement–soil pipe piles. The model’s validity was confirmed by triaxial tests and on-site single-pile bearing assessments.
• The diameter of cement–soil pipe piles significantly influences bearing capacity, with an increase from 600 mm to 1000 mm enhancing capacity by 2.14 times, while wall thickness has a minimal impact. The typical effective length is around 10 m. Larger diameters improve overall capacity but decrease per unit volume. A 600 mm diameter and 150 mm wall thickness are optimal for efficient material use.
• This study explores how the dimensions of cement–soil pipe piles affect their load-bearing capacity. Future research will address factors like lateral friction, end resistance, axial load, and soil core influence on bearing properties, enhancing understanding of load-bearing mechanisms.

7.2. Prospects

Due to the limited time and energy of the authors, this study has certain limitations; the laboratory experiment has not been damaged, so this research can be carried out here in the future, and other materials can be tried for research.