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STADYN Wave Equation Program 7: Revision of Soil Properties (Results)

Posted on 25 June 201826 June 2018 by Don Warrington

In the last post we discussed the change in cohesion in the $\xi-\eta$ interpolation. Now we apply these to the program to see how they change the results.

In the original study there are three test cases, as noted here. For this stage the first and the third will be considered. The results for the second (after adjustments for changes in the $\xi-\eta$ interpolation) are little different from before; we will consider major changes to impact that case in our next round.

That leaves the first and third. For the first–comparison with a static load test–it was necessary to readjust the values for $\eta$ to reflect changes in the meaning of $\eta$ relative to soil properties, as discussed in the last post. The changes made resulted in layering with the following characteristics:

Layer	Bottom y-coordinate, m	$\xi$	$\eta$
1	5.18	-1	0
2	7.32	-1	0
3	15.2	0.5	0
4	30.5	0.5	0

The static load test results can be seen below.

Comparison with the original study show little improvement in the failure load correlation; however, the load-deflection relationship before failure is significantly improved. The system was much softer in the original study and the improvement reflects the better estimate of the elastic modulus. The blow count is very similar to the original study.

Turning to the inverse case that was also discussed in Warrington and Newman (2018), based on the previous results, it was decided to drop consideration of 1-norm results. The results are reproduced in outline below for the three cases run (refer to Warrington and Newman (2018) for details.)

$\eta$ Limiting	Difference Sum	Static Load, kN	Average Shaft $\xi$	Average Shaft $\eta$	Toe $\xi$	Toe $\eta$
+-1	0.0034	673	0.084395	-0.46825	-0.992	0.758
+-2	0.00327	449	-0.455	-1.3375	-0.57	-0.727
+-3	0.00143	269	-0.26775	-2.1775	0.113	1.12

Comments:

The difference sum for the highest $\eta$ variation was the best match we have obtained to date for this case; the velocity match is shown below. The situation around L/c = 1.5 is still difficult but the rest of the correlation is improved.
The static load decreases with broader $\eta$ variation and a lower difference sum. This is different than Warrington and Newman (2018), where the static load “settled down” to a close range of values.
The plot above does not reflect that, on the whole, the variation of $\eta$ along the shaft was less than experienced in the past, especially with the +-3 run. The stratigraphy of the site suggested relatively uniform soil properties along the shaft, and this is beginning to be seen in the inverse results.
The +- 3 run was very heavily “toe weighted” in resistance.

While both of these cases show progress, the time has come to consider the whole issue of pile-soil interface issues, which will be considered in our next updates.

STADYN Wave Equation Program 6: Revision of Soil Properties (Cohesion)

Posted on 24 June 201824 June 2018 by Don Warrington

In our last post we discussed the overhaul of STADYN’s $\xi-\eta$ system relative to the modulus of elasticity, which additionally involved revising the way the program estimated dry unit weight and void ratio. The last is necessary because the modulus of elasticity is estimated using the Hardin and Black formulation. In this post we will discuss revision of another parameter, namely soil cohesion.

We based the relationship of $\rho$ to $\eta$ based on work for the TAMWAVE program. It would doubtless be useful to state the relationship between $\eta$ and the consistency/density of the soil, and this is as follows:

$\eta$	Cohesive Designation	Cohesionless Designation
-1	Very Soft	Very Loose
-0.6	Soft	Loose
-0.2	Medium	Medium
0.2	Stiff	Dense
0.6	Very Stiff	Dense
1	Hard	Very Dense

Doing it this way enabled us to have a linear relationship between $\rho$ and $\eta$ . It is too much to expect for the linear relationship to extend to other variables, and this is certainly the case with cohesion. Unfortunately, a conventional $\xi-\eta$ interpolation dictates such a relationship. The original $\xi-\eta$ function for cohesion can be seen below, for values of cohesion in kPa.

Note that the relationship between cohesion and $\eta$ is linear for the purely cohesionless state at $\xi = 1$ . If extended past the bounds of the graph for lower values of $\eta$ , the cohesion becomes negative. STADYN prevents this from happening but this essentially deprives soft soils of any cohesion.

Baseon on the TAMWAVE values, for purely cohesive soils the following approximate relationship can be established for cohesion:

$\frac{c}{p_{atm}} = 0.5e^{1.5\eta},\,\xi=1$

where $c$ is the soil cohesion and $p_{atm}$ is the atmospheric pressure. The left hand side of the equation is the “normalised” cohesion using the atmospheric pressure. Doing this for parameters such as effective stress makes for an interesting look at soil properties. The best known use of this is in the SPT correction for overburden.

For cases where $\xi < 1$ , the value can be reduced linearly so that $c = 0$ when $\xi = -1$ . The result of all this can be seen in the graph below.

The curve “flattens out” for lower values of $\eta$ , so preventing negative values of cohesion is unnecessary.

In our next post we will look at the results when this is applied to the STADYN program.

STADYN

STADYN Wave Equation Program 5: Revision of Soil Properties (Modulus of Elasticity)

Posted on 23 June 201823 June 2018 by Don Warrington

This is the first post since the presentation of Warrington and Newman (2018) and it introduces what is probably the most significant revision since the original study: the values of the modulus of elasticity for various values of the dimensionless parameters $\xi$ and $\eta$ .

Quick Overview of the “xi-eta” Concept

The basics of the concept are shown below.

The mathematical implementation is shown here:

The results for this (in this case, for the modulus of elasticity) can be seen here:

The original $\xi-\eta$ value matrix was never intended to be “set in stone.” We have already made one revision, by using Jaky’s Equation to estimate the at-rest lateral earth pressure coefficient, and thus Poisson’s Ratio $\nu$ .

The results for this are outlined in Warrington and Newman (2018).

Introduction of Hardin and Black Method for Computing Elastic Modulus

Probably the least satisfactory estimate from the original $\xi-\eta$ scheme was the modulus of elasticity. It was improving this correlation that has led to most of the changes implemented. The research relative to the TAMWAVE project was used extensively in this revision.

First, the Hardin and Black method for estimated the shear modulus of a soil is described here. It allows estimating the small-displacement shear modulus knowing the void ratio, effective stress and the lateral earth pressure coefficient. The last parameter is described above; the void ratio is a function of the specific gravity, the dry unit weight and the saturation state of the soil. If we add the stratigraphic configuration of the soil, we can estimate the effective stress. We can thus estimate the shear modulus–and by extension the elastic modulus–of the soil using some very basic soil properties.

We first turn to the dry unit weight. The TAMWAVE project indicated some revision of how this was computed was in order. It also indicated that, with some generalisation, the dry unit weight could be stated as strictly a function of $\eta$ from all soils. Doing this is a good step to simplifying the scheme and thus improving our chances for success in the optimisation process.

Based on the data, the following correlation was proposed for dry density in $\frac{kg}{m^3}$ :

The dry density essentially varies from $1200 \frac{kg}{m^3}$ at $\eta = -1$ to $2000 \frac{kg}{m^3}$ at $\eta = 1$ .

The specific gravity is unchanged from the original formulation, and is as follows:

Combining the two (we apologise for the mixed graphic formats) yields the following for void ratio:

Because of the nature of the Hardin and Black formula (and of the soils themselves,) the program has an upper limit for the void ratio of 2 and a lower one of 0. This replaces the previous limitations of modulus of elasticity and density.

At this point we need to consider the effect of hyperbolic softening on the shear and elastic moduli. This topic was discussed here and, as was the case with TAMWAVE, we selected a ratio of 0.15. Applying that (and assuming an effective stress of 1 atmosphere) yields the following estimate for elastic moduli in MPa:

In addition to having a stronger basis of fact in the soil properties, by limiting the void ratio the possibility that the correlation goes negative is also avoided.

Soil Mechanics

More Uses for p-q Diagrams

Posted on 6 June 20186 June 2018 by Don Warrington

In our last post on p-q diagrams we discussed their basic concept and application. In this post we’ll expand on that for two applications: using it to estimate the friction angle and cohesion for multiple triaxial tests, and using it to plot the failure function.

Processing Triaxial Test Results

The process of determining internal friction angle and cohesion from successive triaxial tests (i.e., those where the confining stress is successively increased) is well known. In the case of two tests, using the standard $\sigma - \tau$ diagram, the tangent line between the two circles is unique (well, there are two of them, but the slopes and intercepts have opposite signs) as shown below.

Stress Failure Envelope (from Verruijt and van Bars (2007))

If we use the p-q diagram, as we saw earlier, the process is even simpler, as two points have a unique line between them.

But what happens with three tests? Mathematically there is no guarantee of a unique line, and given the nature of geotechnical testing it is the extraordinary lab which could hit such as result. It’s also possible that the failure envelope is non-linear, as shown below.

So is there a way to at least get a decent approximation without guesswork or graphics skills? The answer is “yes” and it involves using p-q diagrams in conjunction with a spreadsheet. The mathematical concept behind this is here and we have an example to show how it is done. The problem is taken from Tchebotarioff’s (1951) classic soil mechanics test. The results of three triaxial tests are as follows, the failure stresses are in tsf:

Test	$\sigma_3$	$\sigma_1$	p	q
1	0.2	0.82	0.51	0.31
2	0.4	1.6	1	0.6
3	0.6	2.44	1.52	0.92

We’ve taken the liberty of computing the p and q values for each test. Now we can plot these in our spreadsheet.

We’ve also taken the liberty to use the spreadsheet’s “trend line” feature to plot a linear “curve fit” for the points. The slope of the equation $m=tan \delta = sin \phi = 0.6041$ , which yields both $\delta = 31.1^\circ$ and $\phi = 37.2^\circ$ . For the intercept $b = c\sqrt {1-\left (\tan(\delta)\right )^{2}} = 0.0001$ , which means we can solve for the cohesion, but in this case the quantity is so small it’s probably best to assume that the cohesion is zero.

The $R^2$ value for this problem is very high, so the correlation is good. We can use this parameter to determine whether we have a good correlation or not. We can also use least-squares trend line analysis for non-linear failure envelopes, although when we consider the “kink” caused by preconsolidation this may not be as meaningful as one would like.

Plotting the Failure Function

As mentioned earlier, the Mohr-Coulomb failure function is define in this way:

A little math transforms this into

$f=2\,q-2\,c\sqrt {1-\left (\tan(\delta)\right )^{2}}-2\,p\tan(\delta)$

$f=2\,q-2\,c\sqrt {1-\left (\sin(\phi)\right )^{2}}-2\,p\sin(\phi)$

Since $\delta$ and $c$ are known, this suggests that we can plot the failure function three-dimensionally. Consider the case where $\delta = \frac{\pi}{8}$ and $c = 5$ . The p-q diagram for the failure envelope $f = 0$ is shown below.

If we plot the failure function three-dimensionally, we obtain this result:

The failure envelope of the previous diagram is the contour line which stops at the q-axis at around q = 4.6. Values below this line are negative and values above it are positive. Positive values of $f$ indicate failure and an illegal stress state. The failure function is used extensively in finite element analyses like this one.

TAMWAVE

TAMWAVE: Cavity Expansion Theory and Soil Set-Up

Posted on 3 May 20183 May 2018 by Don Warrington

One of the things that was attempted in the TAMWAVE project is the use of cavity expansion theory to estimate soil set-up in cohesive soils. Doing this, however, brought some complications that need some explanation.

Cavity expansion theory is basically the study of what happens when one body expands inside of another. When this takes places, additional radial stresses (most analyses center around a cylinder or sphere) are generated. In the case of driven piles, these additional stresses add to the pile’s resistance to load. It can be argued that cavity expansion is one of the key advantages of driven piles. In the case of drilled piles such as drilled shafts or auger-cast piles, this does not take place, as the soil is removed either before or during the actual pile installation. The result of this is that driven piles, for the same length and diameter as a corresponding drilled pile, a driven pile will have a greater resistance to load (ultimate capacity.)

Applying cavity expansion theory to piles has a long history and is detailed in documents such as Randolph, Carter and Wroth (1979) and Yu and Houlsby (1991). Our particular interest is with clays because, in addition to the changes in the soil from cavity expansion, the pore water pressure increases. This is the primary (but not the only) reason why the SRD (soil resistance to driving) in cohesive soils is significantly less than the ultimate capacity; this fact inevitably complicates drivability studies.

The increase in pore water pressure is a dynamic phenomenon; it experiences a sudden increase during driving and then gradually dissipates after installation. How gradually the latter takes place depends on many factors such as the permeability of the soil. Study of this phenomenon is well represented in the literature; however, for the TAMWAVE project it is not really of interest. The primary interest here is the value of SRD after the immediate increase of pore water pressures during driving.

Both cavity expansion theory and practice show that excess pore water pressures can easily exceed the effective stresses. In principle, considering that we have established the beta (effective-stress dependent) method for both cohesionless and cohesive soils, this can mean a complete loss of SRD. Although dramatic drops in SRD are not unknown, for most piles this is unrealistic. The reason for this is that, like the dissipation, the build-up of pore water pressures is a dynamic phenomenon, albeit in a much smaller time frame. Dissipation, hindered as it is by the low permeability of cohesive soils, begins immediately. Pore water pressures (along with any other stresses induced by cavity expansion) also vary with the distance from the pile.

The result of all this is that prediction of both the increase of pore water pressure and its effect on the SRD of the pile during driving is a complicated phenomenon that is not completely represented by closed-form cavity expansion based solutions. For a project such as this, what we need is something that will give a reasonable representation of soil set-up for cohesive soils. To accomplish this, we stick with computing the excess pore water pressure, but with a different methodology. We assume the following:

The basic validity of our effective-stress based beta methods of shaft friction calculation.
All of the decrease from static ultimate capacity to SRD takes place due to pore water pressure increase.
The excess pore water pressure increases affect the effective stress used to compute the shaft friction.
Only those pile segments under the water table are subject to this analysis.

Assuming all that, for a soil set-up factor (from this source, loosely adapted) $S_r$ , the excess pore water pressures that affect the effective stress (which in turn determine the shaft friction) are computed by the equation

$\Delta u = \sigma'_{vo}\left( 1 - \frac{1}{S_r} \right)$

This gives identical results to those when the $S_r$ are applied directly.

In practice this phenomenon is still subject to investigation. Some of the research involves use of numerical methods (such as finite element methods) to simulate cavity expansion effects. This is doubtless an advance over the closed-form solutions of the past, and a necessity given the complexity of the physics of the problem. Empirical methods are also still being developed, such as are documented by Wang, Verma and Steward (2009).