{"id":65,"date":"2026-08-31T23:43:41","date_gmt":"2026-08-31T23:43:41","guid":{"rendered":"https:\/\/tmgmfg.com\/blog\/?p=65"},"modified":"2026-08-31T23:43:41","modified_gmt":"2026-08-31T23:43:41","slug":"boring-log-helical-pile-design","status":"publish","type":"post","link":"https:\/\/tmgmfg.com\/blog\/boring-log-helical-pile-design\/","title":{"rendered":"Reading a Boring Log for Helical Pile Design: A Worked Example"},"content":{"rendered":"\n

Take the bearing stratum depth and N-value off the log, correct N for hammer energy and overburden, convert to a friction angle, compute effective overburden stress at each helix using buoyant unit weights, and sum Ah\u00b7q’\u00b7Nq across the plates. Divide the required ultimate by Kt and you have the installation torque to write on the drawing.<\/strong><\/p>\n\n\n\n

That is the workflow, and it is not just convention \u2014 it is IBC \u00a71810.3.3.1.9, which sets the allowable axial load of a helical pile at Pa = 0.5 Pu (Equation 18-4)<\/strong> and defines Pu as the least of several quantities, the first being the sum of the areas of the helical bearing plates times the ultimate bearing capacity of the soil.<\/em> Florida adopts the model IBC deep-foundation provisions through the Florida Building Code. The factor of safety of 2 you have seen your whole career is that 0.5.<\/p>\n\n\n\n

This post runs the method end to end on a realistic west-central Florida profile. The first configuration we try fails by a factor of two<\/strong>. The second one gets within a percent and a half of passing \u2014 and still fails, on every assumption we test it against. Only the third works. That sequence is the useful part. A clean worked example teaches you the arithmetic; a failed one teaches you where the arithmetic bites.<\/p>\n\n\n\n

We build both the rigs that produce these logs and the piles designed from them, which is the only reason we can write this particular post.<\/p>\n\n\n\n

What a pile designer actually uses<\/h2>\n\n\n\n
\n\n\n\n\n\n\n\n\n\n\n
Column<\/th>\nWhat it’s for<\/th>\n<\/tr>\n<\/thead>\n
Depth and layer boundaries<\/strong><\/td>\nSets helix elevations, defines top and bottom of the bearing stratum<\/td>\n<\/tr>\n
N-value \u2014 all three 6-inch increments<\/strong><\/td>\nSource of \u03c6 and Su. Never use only the summed N; the increments reveal shell and gravel spikes and weight-of-rod zones<\/td>\n<\/tr>\n
USCS symbol and description<\/strong><\/td>\nDecides whether you’re on the drained branch (\u03c6, Nq) or the undrained one (Su, Nc = 9)<\/td>\n<\/tr>\n
Groundwater \u2014 at drilling, at 24 hours, seasonal high<\/strong><\/td>\nEffective overburden. In sand this drives capacity directly<\/td>\n<\/tr>\n
Sample recovery<\/strong><\/td>\nZero recovery is data \u2014 raveling, running sand, or a void<\/td>\n<\/tr>\n
Drilling notes<\/strong><\/td>\nFluid loss and “fell by weight of rod” are karst indicators<\/td>\n<\/tr>\n
Termination depth<\/strong><\/td>\nA hard constraint. You cannot design a helix below the depth you explored<\/td>\n<\/tr>\n<\/tbody>\n<\/table><\/figure>\n\n\n\n

Largely ignored for capacity: moisture content, Atterberg limits except to pick a correlation and flag organics, and the minus-200 fraction beyond classification.<\/p>\n\n\n\n

What makes a stratum worth bearing in<\/h3>\n\n\n\n

Worth separating what the code says from what is judgment, because most published lists blur them:<\/p>\n\n\n\n

\n\n\n\n\n\n\n\n\n\n\n\n\n
Criterion<\/th>\nValue<\/th>\nStatus<\/th>\n<\/tr>\n<\/thead>\n
Allowable load = half the ultimate<\/td>\nPa = 0.5 Pu<\/td>\nCode<\/strong> \u2014 IBC \u00a71810.3.3.1.9, Eq. 18-4<\/td>\n<\/tr>\n
Firm soil for bearing and lateral support<\/td>\nN \u2265 5<\/td>\nCode-adjacent<\/strong> \u2014 AC358 \u00a73.11.2.1 defines firm N\u22655, soft 0<N<5, fluid N=0<\/td>\n<\/tr>\n
Practical minimum for a helix bearing layer<\/td>\nN \u2265 10\u201315 sand, 8\u201310 clay<\/td>\nJudgment<\/strong> \u2014 in no code<\/td>\n<\/tr>\n
Inter-helix spacing<\/td>\n3 \u00d7 diameter of the lower plate<\/td>\nIndustry practice<\/strong> \u2014 stated in the ESRs, not in the IBC. AC358 Table 3 sets 2.4D\u20133.6D as the conformance window for using the default Kt<\/td>\n<\/tr>\n
Pile-to-pile spacing before group effects<\/td>\n3D clear \/ 4D center-to-center<\/td>\nCode-adjacent<\/strong> \u2014 AC358 \u00a76.7 (this is the pile<\/em> spacing rule, often misquoted as the helix rule)<\/td>\n<\/tr>\n
Uppermost helix depth in tension<\/td>\n\u2265 12 \u00d7 largest helix diameter<\/td>\nCode-adjacent<\/strong> \u2014 AC358 \u00a74.4.1.1, \u00a76.9, tension only<\/td>\n<\/tr>\n
Bearing layer thickness below the lead helix<\/td>\n\u2265 3 \u00d7 helix diameter commonly cited<\/td>\nJudgment<\/strong> \u2014 no code number exists<\/td>\n<\/tr>\n
Depth for “deep” bearing behavior<\/td>\n> 5 \u00d7 largest helix diameter<\/td>\nPractice<\/strong><\/td>\n<\/tr>\n
Same bearing layer present in every boring<\/td>\n\u2014<\/td>\nJudgment, and the most violated rule<\/strong><\/td>\n<\/tr>\n<\/tbody>\n<\/table><\/figure>\n\n\n\n

Three borings, one site, three different answers<\/h2>\n\n\n\n

Before any arithmetic, the thing that should worry you most. These are real borings from a City of Tampa project, all on one site:<\/p>\n\n\n\n

\n\n\n\n\n\n
Depth 16\u201320 ft<\/th>\nB-01<\/th>\nB-02<\/th>\nB-03<\/th>\n<\/tr>\n<\/thead>\n
Material<\/td>\nSoft clayey weathered limestone<\/strong><\/td>\nNo recovery<\/strong><\/td>\nSoft<\/strong> sandy clay<\/td>\n<\/tr>\n
N-value<\/td>\n12<\/td>\n5<\/td>\n12, then 4<\/td>\n<\/tr>\n<\/tbody>\n<\/table><\/figure>\n\n\n\n

At the depth where you would want your helix, one boring found weathered limestone, one recovered nothing at all, and one found soft clay that gets softer. A design tuned to B-01 could be off by a factor of three at the pile nearest B-03.<\/p>\n\n\n\n

Another regional example: two adjacent borings where the top of weathered limestone varied from 18.5 ft to 28.5 ft<\/strong> \u2014 a ten-foot swing between neighbors.<\/p>\n\n\n\n

This is the Florida condition, and it is why “we have a boring log” and “we have a design basis” are not the same sentence.<\/p>\n\n\n\n

\"\"
Three borings from one City of Tampa site. At the depth you would want a helix, they disagree completely.<\/figcaption><\/figure>\n\n\n\n

From N-value to soil parameters<\/h2>\n\n\n\n

Correct N first<\/h3>\n\n\n\n

Published correlations are written for N60<\/strong> \u2014 N corrected to 60% hammer energy. Field N is not N60.<\/p>\n\n\n\n

N60 = N \u00d7 (hammer efficiency \/ 60), with adjustments for borehole diameter, sampler, and rod length. Automatic hammers commonly run 75\u201385% efficiency; safety hammers 55\u201360%.<\/p>\n\n\n\n

The correction does not always go up.<\/strong> An automatic hammer at 80% turns a field N of 25 into N60 \u2248 33 \u2014 a 33% increase. A safety hammer at 55% turns that same field N of 25 into N60 \u2248 23 \u2014 a decrease. Which direction you move depends entirely on the hammer, and the hammer is in the report’s methodology section, if it is stated at all. Assuming every correction is upward is how you end up unconservative on a safety-hammer log. (We wrote a whole post on why that number moves so much.)<\/p>\n\n\n\n

For granular soils you then normalize for overburden: CN = \u221a(Pa\/\u03c3’v0), giving (N1)60. Because Florida water tables are shallow, effective stress is low and CN often lands between 1.3 and 1.7<\/strong> at helix depth \u2014 a large upward correction.<\/p>\n\n\n\n

Then convert \u2014 and look at the spread<\/h3>\n\n\n\n

At our example point \u2014 field N = 25, safety hammer taken at 60% so N60 = 25, effective overburden 1,058 psf, giving CN = 1.41 and (N1)60 = 35:<\/p>\n\n\n\n

\n\n\n\n\n\n\n\n\n
Correlation<\/th>\nArgument<\/th>\n\u03c6<\/th>\n<\/tr>\n<\/thead>\n
Peck-Hanson-Thornburn (Wolff 1989 fit)<\/td>\n(N1)60 = 35<\/td>\n37.0\u00b0<\/strong><\/td>\n<\/tr>\n
Peck-Hanson-Thornburn (Wolff 1989 fit)<\/td>\nN60 = 25<\/td>\n34.3\u00b0<\/strong><\/td>\n<\/tr>\n
Industry helical correlation, \u03c6 = 0.28N + 27.4<\/td>\nN = 25<\/td>\n34.4\u00b0<\/strong><\/td>\n<\/tr>\n
Hatanaka & Uchida (1996), \u03c6 = \u221a(20\u00b7(N1)60) + 20<\/td>\n(N1)60 = 35<\/td>\n46.6\u00b0<\/strong><\/td>\n<\/tr>\n
Kulhawy & Mayne (1990)<\/td>\nN60 = 25, \u03c3’v\/Pa = 0.50<\/td>\n46.1\u00b0<\/strong><\/td>\n<\/tr>\n<\/tbody>\n<\/table><\/figure>\n\n\n\n

A 12-degree spread from one N-value.<\/strong> Because Nq is exponential in \u03c6, going from 34\u00b0 to 46.6\u00b0 is a factor of more than five<\/strong> on computed helix capacity. Note also that Hatanaka & Uchida is defined on (N1)60, not N60 \u2014 feeding it raw N60 is a common error that lands you about 4\u00b0 low and quietly looks reasonable.<\/p>\n\n\n\n

Hatanaka & Uchida and Kulhawy & Mayne are known to run high for fine, rounded, uniform quartz sands \u2014 which is exactly what Florida has. Most Florida practitioners cap \u03c6 at 32\u201335\u00b0 for medium dense fine sand regardless of what the correlation returns. That cap is judgment, not derivation<\/strong>, and anyone presenting it as a calculation is overselling it.<\/p>\n\n\n\n

We will design at \u03c6 = 34\u00b0<\/strong>, the value the two conservative correlations agree on. Note what that means: we normalize N for overburden to see the spread, then design off the un-normalized<\/em> branch. That is deliberate and conservative \u2014 using (N1)60 = 35 would give \u03c6 = 37\u00b0 and about 46% more capacity \u2014 but it should be stated, not buried, because a reviewer will ask which one you used.<\/p>\n\n\n\n

For clays, Su \u2248 125\u00b7N psf is the traditional relation. Treat it with real suspicion. Reid and Taylor’s re-analysis of the underlying dataset (Ground Engineering<\/em>, July 2010) found the multiplier ranging from 0.18 to 19.30 kPa per blow<\/strong> with R\u00b2 below 0.2<\/strong> \u2014 no significant association at all. Their mean of about 4 kPa\/blow is roughly 84 psf\/blow, which puts the traditional 125 psf\/blow (\u2248 6 kPa\/blow) about 50% above<\/em> the re-analyzed mean \u2014 the unconservative side. Use SPT-to-Su in soft clay for screening. For design, get a load test or a CPT cross-check.<\/p>\n\n\n\n

The profile<\/h2>\n\n\n\n

A composite that reflects what Tampa-area logs actually show:<\/p>\n\n\n\n

\n\n\n\n\n\n\n\n\n
Depth (ft)<\/th>\nDescription<\/th>\nUSCS<\/th>\nN<\/th>\n\u03b3 moist<\/th>\n\u03b3 sat<\/th>\n<\/tr>\n<\/thead>\n
0\u20136<\/td>\nVery loose to loose fine SAND<\/td>\nSP<\/td>\n4<\/td>\n105<\/td>\n110<\/td>\n<\/tr>\n
6\u201312<\/td>\nLoose clayey fine SAND<\/td>\nSC<\/td>\n7<\/td>\n110<\/td>\n115<\/td>\n<\/tr>\n
12\u201316<\/td>\nMedium dense silty fine SAND<\/td>\nSP-SM<\/td>\n14<\/td>\n112<\/td>\n118<\/td>\n<\/tr>\n
16\u201332<\/td>\nMedium dense fine SAND with shell<\/td>\nSP<\/td>\n25<\/strong><\/td>\n115<\/td>\n122<\/td>\n<\/tr>\n
32+<\/td>\nWeathered LIMESTONE<\/td>\n\u2014<\/td>\n15\u201350+<\/td>\n\u2014<\/td>\n\u2014<\/td>\n<\/tr>\n<\/tbody>\n<\/table><\/figure>\n\n\n\n

Water table measured at 4 ft; seasonal high estimated at 2 ft.<\/strong> Boring terminated at 35 ft.<\/p>\n\n\n\n

We design to the seasonal high, not the measured reading.<\/strong> Every effective stress below is computed with the water table at 2 ft. This costs about 7% of capacity relative to the 4-ft reading, and it is not optional \u2014 a March boring in Tampa is not the condition your pile will see in September.<\/p>\n\n\n\n

Effective overburden, buoyant below the water table:<\/strong><\/p>\n\n\n\n

q'(18 ft) = 2(105) + 4(110\u221262.4) + 6(115\u221262.4) + 4(118\u221262.4) + 2(122\u221262.4)\n          = 1,058 psf\nq'(20)=1,177   q'(21)=1,236   q'(21.5)=1,266   q'(23.5)=1,385\nq'(24.5)=1,445  q'(26.5)=1,564  q'(27)=1,594   q'(29)=1,713 psf\n<\/code><\/pre>\n\n\n\n
Which Nq<\/h3>\n\n\n\n
Two equations are in common use in the helical industry, and neither is more official than the other:<\/p>\n\n\n\n
    \n\n
  • Perko (2009),<\/strong> after Meyerhof: Nq = 0.5(12\u03c6)^(\u03c6\/54) \u2192 22.0<\/strong> at 34\u00b0<\/li>\n\n\n
  • A variant widely used in the industry’s technical literature,<\/strong> which adds a unity term: Nq = 1 + 0.56(12\u03c6)^(\u03c6\/54) \u2192 25.7<\/strong> at 34\u00b0<\/li>\n\n<\/ul>\n\n\n\n
    A 17% swing before any soil variability is considered. We carry both through every calculation below.<\/strong> A design that only works on the friendlier of two equally citable equations is not a design.<\/p>\n\n\n\n
    Attempt one \u2014 short by a factor of two<\/h2>\n\n\n\n
    Design load 40 kip compression. Pa = 0.5Pu, so we need 80 kip ultimate.<\/strong><\/p>\n\n\n\n
    Try a 2\u215e-inch pipe shaft with a 10-inch lead at 21 ft and a 12-inch upper at 18 ft \u2014 3.0 ft apart, which clears 3D on the lower<\/em> (10-inch) plate, 2.5 ft. Net projected areas, plate minus shaft: 12 in = 0.740 ft\u00b2, 10 in = 0.500 ft\u00b2.<\/p>\n\n\n\n
    With Nq = 25.7:<\/strong><\/p>\n\n\n\n
    12 in @ 18.0 ft : 0.740 \u00d7 1,058 \u00d7 25.7 = 20,122 lb\n10 in @ 21.0 ft : 0.500 \u00d7 1,236 \u00d7 25.7 = 15,898 lb\n                                  Qult = 36,020 lb = 36.0 kip\n                                Qallow = 18.0 kip\n<\/code><\/pre>\n\n\n\n
    With Nq = 22.0:<\/strong> Qult = 30.8 kip, Qallow = 15.4 kip.<\/p>\n\n\n\n
    We needed 80 kip ultimate. We have 31 to 36.<\/strong> Short by more than a factor of two.<\/p>\n\n\n\n
    Run it backwards to see how far off:<\/p>\n\n\n\n
    Required average unit bearing = 80,000 \/ 1.241 ft\u00b2 = 64,482 psf\nArea-weighted average q' across the two helices = 1,130 psf\nRequired Nq = 64,482 \/ 1,130 = 57.1   \u2192  \u03c6 \u2248 40\u201341\u00b0\n<\/code><\/pre>\n\n\n\n
    N = 25 does not produce \u03c6 = 41\u00b0 in Florida fine sand under any correlation a reviewer would accept.<\/strong><\/p>\n\n\n\n
    It fails a second time, independently<\/h3>\n\n\n\n
    T required = 80,000 \/ 9 = 8,889 ft-lb\n<\/code><\/pre>\n\n\n\n
    Kt = 9 ft\u207b\u00b9 for 2.875-inch round shafts, per AC358 \u00a73.13.1.1.<\/p>\n\n\n\n
    Torque ratings for 2\u215e-inch round shafts published in ICC-ES evaluation reports and manufacturer data run from about 5,500 ft-lb<\/strong> through 6,400, 7,900, and 8,000 to 8,200 ft-lb<\/strong>, varying with wall thickness and coupling type, with one heavier-wall product rated 11,000. Most of the common ones sit below<\/em> 8,889.<\/p>\n\n\n\n
    Take a shaft rated 8,000 ft-lb: 8,000 \u00d7 9 = 72 kip ultimate, 36 kip allowable \u2014 below the 40 kip design load before soil is considered at all.<\/strong><\/p>\n\n\n\n
    Two independent checks, both failing. That is the useful signal. When the bearing calculation and the torque calculation disagree with your load, they are usually agreeing with each other.<\/p>\n\n\n\n
    Note what they actually say here: the computed soil capacity of 36.0 kip implies 36,000 \/ 9 = 4,000 ft-lb<\/strong> of installation torque, which is about what a 10\/12 on 2\u215e-inch pipe really reads in medium dense Florida sand. The methods agree. The load is the problem.<\/p>\n\n\n\n
    (One honesty note on that cross-check: the bearing sum counts only the helices, while the torque correlation predicts total pile capacity including shaft friction over 21 feet of soil. They are not the same quantity. The bearing sum is the conservative one, and the agreement here is close enough to be informative, not close enough to be a proof.)<\/em><\/p>\n\n\n\n
    Attempt two \u2014 one and a half percent short, which is still short<\/h2>\n\n\n\n
    Go up a shaft size and add a plate. 3\u00bd-inch pipe, triple helix 10\/12\/14<\/strong>: 10-inch lead at 27 ft, 12-inch at 24.5 ft, 14-inch at 21.5 ft. Spacings are 2.5 ft and 3.0 ft, which is 3D off each lower plate, and this is a real manufactured lead-section geometry.<\/p>\n\n\n\n
                                            Nq = 25.7        Nq = 22.0\n10 in @ 27.0 ft : 0.479 \u00d7 1,594 \u00d7  \u2192     19,606 lb        16,784 lb\n12 in @ 24.5 ft : 0.719 \u00d7 1,445 \u00d7  \u2192     26,686 lb        22,844 lb\n14 in @ 21.5 ft : 1.002 \u00d7 1,266 \u00d7  \u2192     32,613 lb        27,918 lb\n                             Qult  =     78.9 kip         67.5 kip\n<\/code><\/pre>\n\n\n\n
    78.9 against 80 needed.<\/strong> On the more favorable of the two Nq equations, at no critical-depth cap, this configuration misses by 1.4%. On the other equation it misses by 16%.<\/p>\n\n\n\n
    78.9 rounds to 79, 79 is “basically 80,” and the temptation to call it close enough is real \u2014 particularly when the number came out of a spreadsheet. A design that fails the check is a design that failed the check.<\/strong> And this one fails on every branch we can test: two Nq equations times three critical-depth conventions is six defensible ways to run the same configuration, and it misses all six, by 1.4% at best and 29% at worst. Go bigger.<\/p>\n\n\n\n
    Attempt three \u2014 a configuration that works<\/h2>\n\n\n\n
    3\u00bd-inch pipe, four helices 10\/12\/14\/14<\/strong>: 10-inch lead at 29 ft, 12-inch at 26.5 ft, 14-inch at 23.5 ft, 14-inch at 20 ft. Spacings of 2.5, 3.0, and 3.5 ft \u2014 3D off each lower plate, all within AC358 Table 3’s 2.4D\u20133.6D window for using the default Kt.<\/p>\n\n\n\n
                                            Nq = 25.7        Nq = 22.0\n10 in @ 29.0 ft : 0.479 \u00d7 1,713 \u00d7  \u2192     21,072 lb        18,039 lb\n12 in @ 26.5 ft : 0.719 \u00d7 1,564 \u00d7  \u2192     28,887 lb        24,728 lb\n14 in @ 23.5 ft : 1.002 \u00d7 1,385 \u00d7  \u2192     35,683 lb        30,546 lb\n14 in @ 20.0 ft : 1.002 \u00d7 1,177 \u00d7  \u2192     30,310 lb        25,947 lb\n                             Qult  =     116.0 kip        99.3 kip\n                           Qallow  =     58.0 kip         49.6 kip\n<\/code><\/pre>\n\n\n\n
    Now the same six-way check, side by side with Attempt two:<\/p>\n\n\n\n
    \n\n\n\n\n\n\n\n\n\n
    Nq<\/th>\nCritical-depth cap<\/th>\nAttempt 2<\/th>\nAttempt 3<\/th>\n<\/tr>\n<\/thead>\n
    25.7<\/td>\nnone<\/td>\n78.9 \u2717<\/td>\n116.0 \u2713<\/strong><\/td>\n<\/tr>\n
    25.7<\/td>\n20D = 23.3 ft<\/td>\n74.9 \u2717<\/td>\n108.0 \u2713<\/td>\n<\/tr>\n
    25.7<\/td>\nflat 20 ft<\/td>\n66.5 \u2717<\/td>\n96.8 \u2713<\/td>\n<\/tr>\n
    22.0<\/td>\nnone<\/td>\n67.5 \u2717<\/td>\n99.3 \u2713<\/td>\n<\/tr>\n
    22.0<\/td>\n20D = 23.3 ft<\/td>\n64.1 \u2717<\/td>\n92.5 \u2713<\/td>\n<\/tr>\n
    22.0<\/td>\nflat 20 ft<\/td>\n56.9 \u2717<\/td>\n82.9 \u2713<\/strong><\/td>\n<\/tr>\n<\/tbody>\n<\/table><\/figure>\n\n\n\n
    Zero for six, then six for six.<\/strong> Attempt three’s worst branch \u2014 conservative Nq, seasonal-high water table, and the most aggressive critical-depth cap anyone applies \u2014 still clears 80 kip. That is what a design you can defend in a plan review looks like: not a number that passes, a number that passes however the reviewer chooses to run it.<\/em><\/p>\n\n\n\n
    Other checks: bearing layer must extend at least 3D below the lead helix. On the lead (10-inch) plate that is 29 + 2.5 = 31.5 ft against a layer bottom at 32 ft \u2014 it clears, barely. On the largest<\/em> (14-inch) plate it would be 32.5 ft, and it would not. The convention is not settled, so say which one you used, and treat 29 ft as the deepest lead elevation this profile supports.<\/p>\n\n\n\n
    Installation torque<\/h3>\n\n\n\n
    T min = 80,000 \/ 7 = 11,429 ft-lb   \u2192  specify 11,500 ft-lb\n<\/code><\/pre>\n\n\n\n
    Kt = 7 ft\u207b\u00b9 for 3.5-inch round shafts, per AC358 \u00a73.13.1.1. The bigger shaft has the lower<\/em> correlation factor<\/strong>, which is why required torque goes up 29% while the shaft only went up 22% in diameter. This surprises people every time.<\/p>\n\n\n\n
    Now the shaft check, and it is not the one you expect. Published 3\u00bd-inch round shaft ratings cluster at 11,000, 13,000, 14,144, and 17,500 ft-lb<\/strong> across the current evaluation reports.<\/p>\n\n\n\n
    A 13,000 ft-lb shaft clears the 11,500 specification with 13% margin, so it looks fine. It is not. If the soil actually delivers the 99 kip our conservative branch predicts, the torque at 29 ft will read 99,300 \/ 7 \u2248 14,200 ft-lb<\/strong> \u2014 and more than that once shaft friction is counted. The installer hits the shaft’s torsional limit and refuses above<\/em> design depth, and now you are on the phone arguing about whether a pile that stopped at 26 ft is acceptable.<\/p>\n\n\n\n
    Specify the 17,500 ft-lb shaft.<\/strong> The rule: your shaft rating has to cover the torque the soil<\/em> will generate at your specified depth, not just the torque your capacity calculation requires.<\/p>\n\n\n\n
    \"\"
    The three configurations side by side, against the 80 kip ultimate the design load requires.<\/figcaption><\/figure>\n\n\n\n
    \"\"
    Two Nq equations by three critical-depth conventions. Attempt two misses all six; attempt three clears all six.<\/figcaption><\/figure>\n\n\n\n
    The critical-depth caveat, stated honestly<\/h2>\n\n\n\n
    Some offices cap effective overburden below a critical depth<\/strong>, on the reasoning that q’ stops increasing linearly in sand. The number matters enormously, and the commonly-repeated “20 feet” is not what the source says. The paper most often cited for it recommends 20D to 30D<\/strong> where D is the largest helix plate diameter, noting published values range from 10D to 40D.<\/p>\n\n\n\n
    For our 14-inch plate that is 23.3 to 35 ft \u2014 meaning at 30D no cap applies to this pile at all, while the flat 20-foot cap in our table is 17D, below<\/em> the low end of the recommended range. We included it anyway, as the most punitive assumption available. State which convention you used.<\/strong> It is judgment, not derivation, and on a marginal design it decides the answer.<\/p>\n\n\n\n
    Writing the acceptance criterion<\/h2>\n\n\n\n
    The output of all this is two numbers on a drawing:<\/p>\n\n\n\n
    <\/blockquote>\n\n\n\n
    Install to a minimum effective installation torque of 11,500 ft-lb, measured as the average over the final 3 feet of advance, and to a minimum lead-helix depth of 29 ft below existing grade.<\/em><\/p>\n<\/blockquote>\n\n\n\n
    Both conditions, not either.<\/strong> Torque alone is not enough \u2014 a pile can hit design torque in a two-foot dense crust at 8 ft and be nowhere near the bearing stratum. Depth alone is not enough either, because the stratum moves between borings.<\/p>\n\n\n\n
    Note the wording: lead-helix depth<\/strong>, not tip elevation. The lead helix on a manufactured lead section sits a few inches above the tip, so a “tip at 29 ft” instruction puts your bearing plate shallower than you designed it. And if you do write an elevation, name the datum.<\/p>\n\n\n\n
    Where this goes wrong<\/h2>\n\n\n\n
    Each of these has a direction. Knowing which errors are safe and which are dangerous matters more than knowing the list.<\/p>\n\n\n\n
    Forgetting buoyancy \u2014 dangerous.<\/strong> Using total instead of effective unit weight below the water table in this profile takes q'(18) from 1,058 to 2,056 psf, nearly double, and Attempt one’s Qult from 36.0 to 70.3 kip. A 95% overestimate<\/strong>, in the direction that gets a foundation built on air.<\/p>\n\n\n\n
    Designing to the measured water table instead of the seasonal high \u2014 dangerous.<\/strong> In this profile it is about 7%. In a flatter one it is more.<\/p>\n\n\n\n
    Using uncorrected N-values \u2014 direction depends on the hammer.<\/strong> On an automatic hammer, raw N under<\/em>-predicts and you leave capacity on the table \u2014 costly, not unsafe. On a safety hammer the correction runs the other way, and treating raw N as N60 is unconservative. Find the energy ratio before you decide which mistake you are making.<\/p>\n\n\n\n
    Feeding N60 into a correlation written for (N1)60 \u2014 conservative here, but wrong.<\/strong> Hatanaka & Uchida on N60 instead of (N1)60 gives about 42\u00b0 instead of 47\u00b0 in this profile. Safe direction, still an error, and it will be caught.<\/p>\n\n\n\n
    A bearing layer that’s too thin \u2014 dangerous.<\/strong> If a helix sits an inch above a change from dense to soft and you only compute at that elevation, the answer is derived entirely from the dense layer. Compute at the helix depth and<\/em> one and two diameters below, and take the lowest. In Florida this bites specifically at the sand-over-weathered-limestone contact \u2014 where, as Tampa B-01 shows, “soft clayey weathered limestone, N=12” can be weaker than the sand above it.<\/strong><\/p>\n\n\n\n
    Ignoring the shaft’s torque rating \u2014 dangerous in the field, not on paper.<\/strong> It does not make the pile weaker; it makes the pile stop short, which is worse because it happens at 4 p.m. with a crew standing around.<\/p>\n\n\n\n
    Skipping the buckling check \u2014 dangerous.<\/strong> With 6 to 12 ft of N = 4 to 7 over the bearing layer, this profile is exactly the case that needs one.<\/p>\n\n\n\n
    Designing below the boring \u2014 dangerous.<\/strong> Those Tampa borings stopped at 20 to 24 ft. Anything specified deeper is extrapolation, and our Attempt three at 29 ft would need a deeper boring than two of those three.<\/p>\n\n\n\n
    Trusting one boring \u2014 dangerous.<\/strong> See the three-boring table above.<\/p>\n\n\n\n
    What the log will not tell you<\/h2>\n\n\n\n
    Corrosion parameters.<\/strong> A standard geotechnical boring log for a building foundation contains none<\/strong> of the electrochemical data a helical design needs: resistivity, pH, sulfates, chlorides, organic content. Those are separate tests that have to be requested \u2014 and on a coastal Florida site they are not optional.<\/p>\n\n\n\n
    Note also that thresholds disagree. AC358’s scope exclusions (\u00a71.2.2) sit at resistivity below 1,000 ohm-cm, pH below 5.5, and sulfates above 1,000 ppm. The FHWA and AASHTO “non-aggressive” criteria are far stricter \u2014 3,000 ohm-cm and 200 ppm. AC358’s numbers are a scope exclusion, not a design criterion<\/strong>, and citing them as though they were a pass mark is a common error.<\/p>\n\n\n\n
    Also absent:<\/strong> the hammer energy ratio, which moves \u03c6 by 5\u00b0 or more. The seasonal high water table, often only estimated. The extent of any karst \u2014 SPT gives you circulation loss and weight-of-rod zones as indicators<\/em>, not geometry. Whether the surficial soil is fill, and how old. And lateral variability, which is the whole point of the three-boring table.<\/p>\n\n\n\n
    \n\n\n\n
    Want a second set of eyes on a configuration?<\/strong> TMG manufactures helical piles, pile caps, brackets, and underpinning products in Tampa, and builds the SPT and CPT rigs that produce the logs behind them. Our Helical Pier Load Calculator<\/strong> is a quick way to check a configuration before it reaches a drawing. Call (813) 464-2299<\/strong>, toll-free 1-888-508-RIGS<\/strong>, or email info@tmgmfg.com<\/strong>.<\/p>\n\n\n\n
    Ramzy Moumneh, TMG Manufacturing \u2014 Tampa, Florida. TMG builds geotechnical drill rigs and deep foundation products.<\/em><\/p>\n\n\n\n
    \n\n\n\n
    FAQ<\/h2>\n\n\n\n
    How do you size a helical pile from a boring log?<\/strong>\nPick a bearing stratum with adequate N-value and thickness, compute effective overburden stress at each proposed helix depth using buoyant unit weights below the seasonal high water table, convert corrected N to a friction angle, and sum Ah(c\u00b7Nc + q’\u00b7Nq) across the plates. IBC Equation 18-4 then sets the allowable load at half that ultimate.<\/p>\n\n\n\n
    What N-value do you need for a helical pile bearing stratum?<\/strong>\nThere is no code number. AC358 defines firm soil as N \u2265 5, which is a lateral-support threshold rather than a bearing criterion. Practical judgment puts a usable bearing layer at N \u2265 10 to 15 in sand and 8 to 10 in clay, with at least three helix diameters of that material below the lead plate.<\/p>\n\n\n\n
    Do you use raw N-values or corrected ones for helical pile design?<\/strong>\nCorrected. Published correlations are written for N60. The correction can go either way \u2014 an automatic hammer pushes N up by about a third, a safety hammer pulls it down \u2014 and because Nq is exponential in friction angle, either error moves computed capacity substantially.<\/p>\n\n\n\n
    How do you calculate the installation torque to specify?<\/strong>\nDivide the required ultimate capacity by the shaft’s Kt factor. AC358 \u00a73.13.1.1 gives Kt = 9 ft\u207b\u00b9 for 2\u215e-inch round shafts and 7 ft\u207b\u00b9 for 3\u00bd-inch, so 80 kip ultimate on a 3\u00bd-inch shaft requires about 11,400 ft-lb. Then check that number against the shaft’s published torsional rating \u2014 and against the torque the soil will actually generate at your design depth, which is usually higher.<\/p>\n\n\n\n
    Why does the water table matter so much for helical piles in Florida?<\/strong>\nHelix bearing capacity in sand is directly proportional to effective overburden stress. Below the water table, buoyancy roughly halves the effective unit weight. In a typical Tampa profile, using total stress instead of effective inflates computed capacity by about 95%.<\/p>\n\n\n\n
    \n\n\n\n
    Sources<\/h2>\n\n\n\n
      \n\n
    • IBC \u00a71810.3.3.1.9 and Equation 18-4<\/strong> (allowable axial load of helical piles); adopted in Florida through the FBC deep-foundation provisions<\/li>\n\n\n
    • ICC-ES AC358<\/strong>, Acceptance Criteria for Helical Pile Systems and Devices<\/em> \u2014 \u00a71.2.2 (corrosion scope), \u00a73.11.2.1 (soil definitions), \u00a73.13.1.1 (Kt values), \u00a74.4.1.1 and \u00a76.9 (tension embedment), \u00a76.7 (pile-to-pile spacing), Table 3 (helix spacing conformance window for default Kt)<\/li>\n\n\n
    • ICC-ES evaluation reports ESR-1854, ESR-3074, ESR-3418<\/strong> and ESR-3982<\/strong>, and current manufacturer data sheets, for the published shaft torque ratings<\/li>\n\n\n
    • Perko, Helical Piles: A Practical Guide to Design and Installation<\/em> (Wiley, 2009) \u2014 Nq = 0.5(12\u03c6)^(\u03c6\/54), after Meyerhof (1976)<\/li>\n\n\n
    • Published industry technical literature on bearing capacity factors for helical pile design (Nq = 1 + 0.56(12\u03c6)^(\u03c6\/54)) and on critical depth in sands (20D\u201330D of the largest plate)<\/li>\n\n\n
    • DFI Helical Pile Foundation Design Guide<\/em>, 1st Edition (2019)<\/li>\n\n\n
    • Peck, Hanson & Thornburn (Wolff 1989 fit); Kulhawy & Mayne (1990); Hatanaka & Uchida (1996)<\/li>\n\n\n
    • Reid, A. and Taylor, J., “The misuse of SPTs in fine soils and the implications of Eurocode 7,” Ground Engineering<\/em>, July 2010<\/li>\n\n\n
    • City of Tampa 30th Street Outfall and Spring Lake Stormsewer geotechnical reports; SWFWMD geotechnical report; FDOT District 7 SR-60 geotechnical memo and Central Florida Sinkhole Evaluation<\/li>\n\n<\/ul>\n","protected":false},"excerpt":{"rendered":"
      Helical pile design from a boring log, worked end to end on a real Tampa profile \u2014 N-values to soil parameters, helix sizing, and the installation torque you specify. 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