## 3 Validation of the Proposed Method
### Analysis Results of a Specific Case Study
#### 3.1.2 Single Bored Pile B7: Analysis Results  
The soil unit weight, \(\gamma\), of the soil was not reported in the article of [PERSON] et al. [43] however it was estimated according to the interpretation of the CPT test data. Along the depth of interest, approximately equal to the first 15 m, corresponding to 10 pile diameters, the mean tip resistance was approximately 5 MPa. The mechanical properties of the pile used in the analysis were the same as those indicated in Table 2. Figure 9 presents the average moment-curvature relationship used in the analysis and computed using the model that considers the tension stiffening. On the basis of the data provided by the CPT tests, a friction angle \(\varphi^{\prime}\) of \(34^{\circ}\) was used, obtained with the correlation proposed by [PERSON] [57]. \[\varphi^{\prime}=17.6+11\cdot\log_{10}\left(\frac{q_{t}-\sigma_{v0}}{\sqrt{ \sigma^{\prime}{}_{v0}P_{atm}}}\right) \tag{12}\]  
The shear modulus profile at small strain levels was that provided by the authors in the SCPT data, but was simplified in the analysis, and thus a \(G_{max}\) profile linearly increasing from 15 to 150 MPa was adopted. The Poisson ratio was taken as equal to 0.35, so \(E_{max}\) linearly increased from 40 to 400 MPa. The non-linear \"p-y\" curve adopted, to model the near-field soil response, was that proposed by [PERSON] et al. [32]. Since the water table was located 1 m below the ground surface suction effects were considered in the analysis. Nevertheless, it was not possible to use the MK-model rigorously due to the lack of information about the actual degree of saturation above the water table. As a consequence, suction was considered by assuming that the upper soil layer was fully saturated by capillarity. This assumption enabled us to use Equation (11) to estimate the vertical effective stresses considering \(\chi=1\) and a linearly increasing suction value from 0 kPa to 10 kPa starting from one-meter depth up to the ground surface. This results in an increase in the ultimate pile-soil interface pressure, computed using the relationships suggested by [PERSON] et al. [32], only in the first meter depth. Figure 10 compares the \begin{table}
\begin{tabular}{c c} \hline \hline
**Pile Diameter \(D\) (mm)** & **1500** \\ \hline Pile Length (m) & 34.9 \\ Cross Sectional area (cm\({}^{2}\)) & 17,672 \\ Concrete compressive strength \(f_{c}\) (MPa) & 27.5 \\ Reinforcement Yield Stress\(f_{y}\) (MPa) & 471 \\ Steel Ratio \(\rho_{s}\) & 0.025 \\ Intact flexural rigidity \(EI\) (GNm\({}^{2}\)) & 6.86 \\ \hline \hline \end{tabular}
\end{table} 
Table 2: Structural properties of bored pile, [PERSON] et al. [43]. Figure 9: Computed ”average moment-curvature” relationship for B7 pile section. measured and computed load-deflection curves. The p-y curves used in [PERSON] et al. [43] follow the shape suggested by [PERSON] [31] but were established using the DMT tests data. The same figure also shows the deflection curve obtained without far-field component. As expected, there is a reduction in computed horizontal displacement. In this case, the reduction varies between 23% and 8% with increasing load level. Pile deflection profiles versus depth are shown (computed and measured) for three different lateral load values at the pile-head, considering suction in Figure 11 and without suction in Figure 12.  
Figure 11: Pile deflections versus depth for B7 single pile under various load levels: Computed data obtained considering suction. Figure 10: Comparison of measured and computed ([PERSON] et al. [43] and proposed method) Lateral Load vs. Head Deflection curve. The agreement is good in both cases however considering suction, there was an improvement in the prediction of the measured data. In addition, Figure 13 compares the bending moments profiles for various load values computed with the proposed method and those obtained by [PERSON] et al. [43] using LPILE. Note that in the work of [PERSON] et al.